MSAR 73-14
AC-860
FINAL REPORT
to
Environmental Protection Agency
Office of Air Programs
Research Triangle Park, North Carolina 277
on
HYDROCARBON POLLUTANT
SYSTEMS STUDY
VOLUME II - APPENDICES
25 January 1973
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ACKNOWLEDGEMENTS
. This study was performed for the Office of Air
Programs of the Environmental Protection Agency under Con-
tract Number EHSD 71-12.
The study was directed by Mr. W.J. Cooper, MSA
Research Corporation program manager. MSAR personnel
contributing significantly to the study were Messrs. W.A.
Everson, J.V. Friel, J.S. Greer and C.A. Palladino.
The MSAR study team was ably assisted by four
subcontractors: Industrial Health Foundation; Patent De-
velopment Associates; Singmaster and Breyer; and University
Science Center. Space does not permit individual acknow-
ledgement of all contributing personnel, but the following
individuals are recognized for their subcontract direction
and contribution: Mr. Harry Bowman, IHF; Dr. B.J. Lerner,
PDA; Mr. Stanley Zukowsky, S&B; and Mr. Edwin Snow, USC.
Mr. VJi11iam R. King, EPA Project Officer provided
assistance. during all phases of literature collection and
review, as well as constructive criticism of the interim
results used in compiling this report.
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TABLE OF CONTENTS
Appendix
Title
A
Calculations, Quotes and Correspondence,.
Gasoline Storage Tanks
B
Calculations, Quotes and Correspondence,
Incineration
D
Calculations, Quotes and Correspondence,
Adsorption
Incineration-Absorption and Scrubbing-
Absorption
C
E
Incineration-Scrubbing Systems for Hydro-
carbon Emission Control
F
Particulate-Scrubbing Systems for Hydro-
carbon Emission Control
G
H
Questionnaire Survey Analysis
Breakdown of Fuel Consumption and Estimated
Emissions and Review of Elementary Combustion
Studies.
I
Waste Combustion Emission Factors and Muni-
cipal Waste Breakdown
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A P PEN D I X "A
". Calculations, Quotes and Correspondence
". Gasoline Storage Tanks
TABLE OF CONTENTS
Control Process Concept
Gasoline Loss Correlations and Calculations
Installed Capital Cost Estimates
Operating Cost Estimates"
Vendor Budget Quotations
Paqe No.
A-l
A-2
A-a
A-17
A-26
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PROCESS CONCEPT FOR CONTROL OF GASOLINE VAPOR EMISSIONS
FROM FIXED ROOF STORAGE TANKS VIA CONVERSION WITH INTERNAL
FLOATING COVERS OR STORING IN PONTOON FLOATING ROOF TANKS
Fixed roof gasoline storage tanks discharge vapors
to the atmosphere by means of tank filling, emptying, and
breathing.
Losses can be considerably reduced by eliminating
the vapor space within the tanks. .
Economic evaluation of pontoon floating roof tanks,
fixed roof tanks, and fixed roof tanks with internal floating
covers is carried out for nominal tank capacities of 50,000,
100,000, and 150,000 barrels.
A-l
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CUSTOMER M~ I~
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REVISION DATE
PAGE NO.' A -, '5
(E 153)
-------
CONSTRUCTION COST ESTIMATE
CUSTOMER Y\ £:, ~
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(E 153)
DATE
REVISION NO.
REVISION DATE
PAGE NO. A - \ ~
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CUSTOMER ""'$ A..
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CONSTRUCTION COST ESTIMATE
CUSTOMER }1S A
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DATE
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REVISION DATE
PAGE No.-A - \ b
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July 13, 1971
Singmaster & Breyer
219 East 44th Street
New York, New York 10017
Attention:
Gentlemen:
Mr. Samuel Martin
Confirming your telephone conversation with our Mr. E.H.
Bodinson on budgetary prices for Cone Roof Tanks and Pon-
toon Floaters, as well as our , we wish
to adivse you of budgetary figures as follows:
Cone Roof Tanks Pontoon Floaters Hammondflotes
In existinC/
Cone Roof
Tanks
95°4> x 40' $108,000 $120,000 $29,000
~20B4> x 48° $188,000 $208,000 $40,000
150'4> x 481 $274,000 $300,000 $56,000
The ~bove prices are basically for the Eastern part of the
United States. It must be u~derstood that those prices,
given above, are simply for budgetary and are not intended
to be use'd as a selling price. .
Very truly yours,
A-26
-------
July 9, 1971
Singmaster & Breyer
235 East 42nd Street
New York, New York 10017
Attention:
Mr. Sam Martin
Storage Tanks
Dear Mr. Martin:
In response to your recent verbal inquiry, we are
attaching a separate tabulation confirming the verbal budget
prices given you for the various sized storage tanks under
consideration.
These prices are for tanks with standard fittings
erected on foundations provided by others, assuming average
erection conditions. We trust that this is adequate for
your needs at this time. Should you require anything further,
please feel free to call on us.
Yours very truly,
A-27
-------
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-------
APPENDIX B
CALCULATIONS, QUOTES AND CORRESPONDENCE
INCINERATION
TABLE OF CONTENTS
CONTROL PROCESS CONCEPT
INCINERATION PROCESS CALCULATIONS
B-1
B-2
INSTALLED CAPITAL COST ESTIMATES
OPERATING COST ESTIMATES
B-22
B-58
VENDOR BUDGET QUOTES
B-94
-------
PROCESS CONCEPT FOR CONTROL OF HEXANE AND BENZENE FUMES
FROM A PAINT BAKING OVEN VIA INCINERATION
A paint baking oven exhausts a SO/50 weight percent
mixture of benzene and hexane vapors in air at concentrations
of 15% and 25% of the lower explosive limit (LEL) and a tem-
perature of 375°F.
Air pollution regulations require that
carbon emissions be reduced by 90% by conversion
dioxide and water vapor before exhausting to the
all hydro-
to carbon
atmosphereo
Economic evaluation of thermal and catalytic incin-
erators both with and without heat recovery systems is per-
formed for oven exhaust volumes of 1,000, 10,000 and 20,000
SCFM. The oven is assumed to operate two shifts per day,
365 days per year.
B-1
-------
NAME OF COMPANV2::1 SA ,R€~>E)..\Q.e,.H
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C.OMI'UTATION SHEET
J. O. No. 'PS - 2.. 2. S
SHEET No. ~..,.2. OF.'
DATE 7/;2.0/7/
SINGl\IASTER (k BREYER
CHEMICAL 80 METALLURGICAL PROCESS ENGINEERS
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CUSTOMER Jv\SA Re~E"'~~\ CO~~
LOCATION~E vA "'$ c.\,y) ,\=>" '
PROJECT YS - 22. S
ACCOUNT
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CUSTOMER.M 5 A. RESEAKC,\-\' C~ ~(>
LOCATION' EVANS, C \1'.'} ?A
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CODE
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DATE
CONSTRUCTION COST ESTIMATE ..
SINGMASTER & BREYER
DESCRIPTION t\7'D~OC.Ioi.t1 C(O~ E""",\5.\ o~
CO~"N.OL.
PROP. NO.
CONT. NO,
C"-"AI..)"T\C- h.1G'I>:\~A"I1'H~W'TI-\ ?\2..\MAoi!.j p.. ~O
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MADE BY
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ESTIMATED COST
CENTS OMITTED
SUB
CONTRACTS
-.
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MATERIALS
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PAGE No._3-4i
(E 153)
-------
CUSTOMER -,"'So A. "RE.~e~~~
LOCATIONJ~:v'A.'~'" C.rr)'1 VA. .
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PROJECT 'PS- 22.8
CoQ.~
ACCOUNT
NUMBER
ITEM & DESCRIPTION
CA.'~\..YTl" '\~Co\\:Ie.\!.A"'Oi.) 'fi,t-\Aq,y ~:
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CONSTRUCTION COST ESTIMATE
SINGMASTER & BREYER
DESCRIPTION HYD~c:..A;!"BQ~ t:..kM\MOr-J$
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PAGE NO.
4q7/Jf~ -
(E 153)
8.46
-------
CONSTRUCTION COST ESTIMATE
CUSTOMER M SA e £'sQ/4?cH
LOCATION EVA/J5 Cfiy. 111
, ,
PROJECT. ,F!;-.:J;J. 8
C04P
SINGMASTER & BREYER
DESCRIPTION fl.yP4~C~;;gO~ E"/'1I$S/~,cl5
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PROP. NO,
CONT, NO,
MAOE BY
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UNIT COSTS ESTIMATED COST
ACCOUNT ITEM & DESCRIPTION QUANTITY UNIT CENTS OMITTEO
NUMBER SUB SUB
LABOR CONTR MAT'L. LABOR CONTRACTS MATERIALS TOTAL
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, -
CONSTRUCTION COST ESTIMATE
~INGMASTER & BREYER
DESCRIPTION flYL'~::Jc~I!E",d
vi A / ""&./~c"12Ailo M
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tlf,Jl rWdW''r rJ~r /?~It()Yo;y MADE BY
APPROVED
ACCOUNT
NUMBER
ITEM & DESCRIPTION
UANTITY UNIT
UNIT COSTS
ESTIMATED COST
CENTS OMITTED
'SUB SUB
LAaOR CONTR MAT'L LABOR CONTRACTS MATERIALS TOTAL
- H02f1I4f.,;..IN(.I"'!~A7"~ :_~I'1.A.e-r~...~_~t:.1f . - .: - . 3+,000.
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CO$7
CODE
-. - -.- .. _. - -- - --
(E '53)
DATE
REVISION NO.
REVISION DATE
PAGE No.B-SO
-------
CONSTRUCTION COST ESTIMATE
CUSTOMER .l1 SA
\2..E$€>\Il.C:H C 0 ~i"
SINGMASTER & BREYER
DESCRIPTION ;-Iyp,Zl1~p",,' 1#1 'rJ-IIlIl" )-/~~ J?GCov£>2Y
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DATE
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-
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-
-
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---
"
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~Qf'c'~Ty~~S'; IX. CAI'IT1<\L. CO$.'\" (.'\U. bOO )
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-------
<. ,...., 0.;;
'I TE "1 D e.~c..1i I~T\ 0 ~ . !ijtR. ~ la" s<..~
~-~........ ..4........... ....,.' ...
RAW MAiE.z, "" LS, 1
UrILITIe.~: "
"
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.-..,
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-_?"y':~~L-t.. O"ct14eA.i)' 30% DI~T t....foIthtZ.- 43g .--......
,_~~N,..OVEr.:H~ SOo/...J).\,z,ecT' LABOR. 730
. -. ~~~-'''('''\\.)<>' ~ ' , ,
" -
- S \-t H' (>\.,)"": -
--
INDIRECT MF6. COST .. 'IlL 9
..: -
'D~I7I.!.GC.'''TI oiJ ~ 10% C"'PITA\.. COS,T l\A4..,~) \ Aq, () .....4
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\..&:~'t.
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. '
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111.4-7
-. 'TOTA\.. f'LANT
LEVEL C.OST' 3S,"2..\
--
c,O~I>OI~"'"t8 -, .~." '.- -........
O""~He:", b> .... '"5% ,.o,..~\.. P.:.Ato)T L&ve.~ COf>r" .. IO~
- .-
,
TOTAL:- MF& "' " 3'b'8~
C 0 <:;T ~
"
...---....-- ~~. -. "
N W\"T\.l~I\\. ~"\S ~\4q,:, <. <;A.~\~M @ CN~~) , L~~s \'1., '\ S' 0
"!,. "'8" 10' ~o( ""~j)" S1:~ "'$'840 ~o/'4.;:aq_''1~o:' 2..~, 1:!1~
"\o4n..'
SA'i, ;z. l,'7Db. 'OOb~
" ~-4"Z.
:
MANl.\t=Ac:n",Q."~c,, E~t"lt1ATt5. OVEN £Y.\4A4~-r COI'-\~...OI- Sc.I-'lSl1e
CO ""TI!..OL.. p(2..0c.e~..) :T'-'IC.I!."'AL \ \.1u~~~....-rIO~ W \'t~ C; ~"t:)QAQ." \.\Ekr' ~~ "" OVe..:! ~~IoIA~~ t.""TIf:
0' . -;- , ~,",I"'~ ,'<',~, \-\R. 10,000 SC:fM
- r>R-Oc.E~ PE.R./~\t\t,)'" ".-'IE. ~ ~tJ"'y" 3 '$" "'IR,. ~1i'4 0 ~ ( .
" . 7.S~ I-E'I..)
-------
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M A~'" "eN'~I~c..", ~
- C;'H~ NER "-I.. ~1 "-\101'10 sJ{ C. J),rI'1"p.t,. G.O!>T (~:s"C,C>Q~) 1\ J\~.s:'o
1'11\'~:", S'H'P.~I €!".-, -- '....
-... - --..
\) I r,l;;c'r l.A.CO~~ . S ~ A.J~"-IIPT'" '30 '''''I j:T/.... cz..... ,-"' ,,,oJ,,,, ~ ,. '. ,+~~
\ 0-% b,c(;.c::r: LA1:."rt ...----.'
-_S\A,p"t:.}\~\C'tJ: -I~ 1----'
,.~:.J \\,~~Ci t\ ~ "(..'. e.C'::.;'\' .' -~,~~~ f-----
- '. . _.~~--
.._P"l'P-~"'l.. O\'':~/-I~Aa 30% '. A.3~
DII7,,~.AG~ ---
,--~~_\.'IJ'\'" OyeRH~L $'0 r.. D I K!;C-r L-~.1?lC!n. -- :or :39.- ---
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CA.",n.i
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f\XE\) M F6 c.osr " .=l. 1."tl
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"
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, ?R.OC.E~ 'OPE.\,;.f.\TIt.)(,. -tV\E.~,'L. ~~~rs. " ~ '.s'O::f:'~f~O *' ( :D, 1>0)0 s~jCl'1
. . .' . , 1...s'~ L.I!L.
-------
Mr. Sam Martin
Singmaster & Breyer
235 East 42nd Street
New York, New York
Dear Mr. Martin:
RECEIVED
JUL 1 G 1971
Singmaster & Breyer, Inc.
July 15, 1971
10017
Confirming our telephone conversationg your design Gonditions
are as follows:
Flow:
Temperature:
Contaminants:
Concentration:
Fuel:
Schedule:
Performance:
Maximum Self-Recup-
erative Exchange:
Correction Equipment:
1000810,000 and 20,000 SCFM
375°F
50/50 mixture hexane and benzene
25% LEL .
Natural gas @ 60~/106 BTU
80% .
Rule 66, 90% conversion
46-47% .
Thermal or Catalytic
As mentioned to you, the 1000 SCFM is too small for, self-recup-
erative equipment and I hesitate to offer more than budget
approximations on catalytic without a Cat 2 program run for
sizing. By copy of this letter to Engineering' with the appro-
priate request form, I am asking for this and will advise you
as soon as it is complete. ' '
Thermally we have the following:
1000 SCFM
Thermal unit only; Model RFA-30 Fumabator; about $9,000; size
4' wide by 4' high by 13' long @ 3200 lbs.; resistance to flow
about 3" w.c. hot; specifications attached.
10,000 SCFM
Thermal unit only; Mode1DT-10,OOO-15.0M-SABH; about, $30,000;
size 8' wide by 8' high by, 24' long @ 10,,000 Ibs.; resistance
to flow 4-5" w.c'. hot; specifications attached.
e. '14-
-------
Mr. Sam Martin
Singmaster & Breyer
July 15, 1971
Page 2
10,000 SCFM, Con't.
Thermal unit with self-recuperative heat exchanger; Model
DTjSR-lO,000-15.0M-SABH, abOut $47,,500; size 8" wide by 8'
high by 34' long',@ 21,000 lbs.; resistance to f.low 10-12" w.c.
hot; specifications attached (same as ,therma:J. unit only specs).
20,000 SCFM
Thermal unit only; Model DT-20,OOO-15.0M-SABH;
size, 10' wide by 10' high by 24' long @ 14pOOO
to flow 4-5w WOCG hot; specifications,attached
for 10,000 unit). .
about $43,000;
lbs.; resistance
(same specs as
Thermal unit with self-recuperative heat exchanger; Model
DTjSR-20,000-15.0M-SABH; about' $75.,500; size 10' .wide by 10'
high by 3.4' long @ 31,000 lbs.; resistance to ,flow 10-12" w.c.
hot; specifications attached (same specs as for 10,000 unit) .
Exhaust temperature on the self-recuperative units is about
900-l000oF. .
I would budget about 50% on a grass roots for installatiol1. and
100% on a retrofit.
Very truly yours,
i».,\(.
-------
August 9, 1971
Mr. S. T. Martin
Singmaster & Breyer
235 East 42 Street
New York, NY 10017
Dear Mr. Martin:
During our recent telephone discussion, I indicated that I would forward
information requested in your letter of July 15. At. that time, I had
not completely reviewed the data and cQnditions stated.
The LEL is hypothetically stated as 25% in air. We' would be normally
quoting only thermal oxidation equipment with a sing~e pass heat exchanger
(45%) in the two larger sizes. The price for these three models would
be approximately $10,000, $50,000 and $70,000. The respective fuel
useage would be as follows: 1.0 MMBTU/HR, 3.0 MMBTU/HR and 6.0 MMBTU/HR.
I .
Many baking applications have contaminant in the range of 5 to 15% LEL.
In this case, catalyst oxidation can be the more practical solution.
In as much as the above does not offer comparative data, you may wish
to contact me for some other conditions of operation.
Very truly yours,
RECEIVED
AU G 111971
. .
Skgmasi~r. & Breyer, Ino..
. ~ "'"
-------
August 13, 1971
Mr. S. T. Martin
Singmaster & Breyer
235 East 42 Street
New York, NY 10017
Dear Mr. Martin:
Enclosed is a revision in our analysis .for the project operating conditions
described in your July 15 letter.
We are now basing our evaluations on 15% LEL. It should be noted that
the catalyst can presently be replaced for approximately $1.00 per
SCFM ($1,000, $10,000, and $20,000 respectively). I would anticipate
15,000 hours of operationas life from the process as desc~ibed ( 50/50
by weight blend Hexane/Benzene). .
Very truly yours,
RECEIVED
AUG 1 G 1971
Singmader & Breyer, .Inc.
.B -,\.,
-------
Singrnaster and Breyer
EPA Contract EHSD-7l-l2
Table 1 -
Thermal Oxidation thnits
SCFM
Without HX
. ,
With HX (45%)
1,000 Fuel
Price
1.0 MMBTU/HR.
$10.,000 .
10.0 r-mBTU/HR
$20,000
4.0 MMBTU/HR
. $50 ~OOO .
8.0 MMBTU,IHR
$70,000 .
10,000 Fuel
Price
20,000 Fuel
. Price
SCFM
Catalytic Oxidation Units
Without HX
With HX (45%)
Table 2
1,000 Fuel
Price
0.4 MMBTU/HR
$11,000
. 4. 0 MMBTU/HR
$30,000
1. 0 MMBTll/HR
$60,000
2'. 0 MMBTU/HR
$90,000
10,000 Fuel
'. Price
20,000 Fuel
Price
Conditions are 15% LEL; 375°F inlet; 1,000°F and'l,400°F.
~ ' I '1 -1 (
r-1C2-
~a)4B- Co.vs~n"';-/tJN F/~",..,.cr
1'f:;'7Z. 7i1.t&'~# - Y /11'/, /
/~- 5~}{
en ,(1:...,,1911. 7',I/I;J ~t),A'Mnc.
F.~. tF{)WIf~D~
("' )
FI'I"') C,,:;.,.
IA"L.'-I"",d
,";/ /J..,Dtnt;1'" P~"..£s
H;:~
~ 4 ""-'
~ ,..5'0
r,:"p;t...
Sr1"L.~E,ff' r~
..,.,~ .:. IE s r
.-
~A.J"
~.'\ 8
-------
, 1~E:Cf~iVCD
SEP131971
, September 10 , 197J
Sifigmaster & Breyer, Ino.
Mr. Sam Martin
Singmaster & Breyer
235 East 42nd Street
New York, New York
2SX L.EI-' CAT~V't'Ic..- COM6L.4.CT'O,)
.ro/.s-o -' H6XII)J~/.Bfl'!J2r;;,.,t
10017
Dear Sam:
Confirming our conversations and furthei to ou.r' letter Qf .July 15,'
this is areal toughie catalytically , particalJ.:( dqa'to the'.. "
solvent composition and concent~ation. In,f~c~, it would be hard
to choose a more difficult combination asevide~cedby the
attached copies of our Cat 5 computer runs, and ~apresents
between 7 and' 8 times the quanity of:catalystne~essary for
easy solvents like 'toluene or xylene, etc. .'
.
Now, summaring the budget costs reported to you"wehave'as
follows:' , '
Flow
, SCFM
W:j..th
, ~xcha'n(ler
'.. " ,',
, Ca't' Urd t
1,000
10,000
,20,000
, $ 15, 000
, $96,500
$178,000
$107 ,~500 '
$2Q3;,000
and you have 'the fuel consumption, 'resistance 'to flow numbers, etc.
In addition, the' incinerator units to handle the solverit recovery
waste gaseous vents specified are, summarized as follows:
, Flow SCFM
, FTanre ,Unit,
1,000
10,000
, 20',000
, $-8 ,000
, "
, $11-,000 -
'$15,000
and only pilot fuel ,is necessary to ignite thexni;xture with excess
air for complete combustion of the solvent lade~ ~team. '
Yours very tr~:ty,
8-'19
-------
. $'i,,~ ~6J'Tt''' r (3, ~ 't,,~
. ," . .
ON AT 11:33:01 .20 JUL 71 TUESDAY
ASSISTANCE: 91~-592-4g51
USER 10, PASSWORD, PROJF.~T ,I D--DAL001, UOPI?OD
SYSTEM--BA$IC
VERSION 14 JUN'71
RF.ADY
LOAD CAT5
RE.4DY'
RUi'~
C5N8f)v' . 137
1?:3~
CAT5
11:35
07/~r)/71
ENTER CONVERSION,C,'; iNLET'TEMPERATURE,Tl, 'DE~~EES F,
AVG GAS FLOW, G, SCFMo
190,600,1000
IF YOU WANT THE SOLVENT C~DE PRINTED, ENTE~ 1; YFNOT,??O
ENTER SOLVENT CODE NUM3ERS. USE ZEROS TO ~AKE e ENTRIES"
114,1,0,0,0,0,0,0 .
ENTEq cnNC~~TRATIONS, Y(N), ATU/SCF.
16.7,6.4~,O,O,0,0,0,O
VSE 0'5 Tn ~AKE 8 ENTRIES
INLET TEM°F.~ATU~E, .TI = 600
OUTLET TEMoERATURE, TO~ 1225,
DELTA T ' = h~5
cm1PONENT CON,CENTQAT J I')N5 AT 5' pos IT IONS HI BED FI')LLO\'/
Y(l). Y(2) Y(3) Y(4) Y(5) .Y(G) y(n 'Y(8) x
--..;-- _"",",CIOII_- ----- ----- ----- ----- ----.- ----wi> -----
J=l 6.6r> 4..17 : ~oo 00 000 000 .1)0 .00 .?3
J=2 6.49 L 8/1 ' .00 .00 .O;i) .00 ..00 .00 . l~ <) .
J=3 5.7f) .22 .00 .00 .00. .00 .!'>O ~OO .q4
J=4 3.68 .00 .00 .00 .00 .00 .00 .00 1.66
J=5 1.3J ~OO ;00 .00 .00 ..00 .00 .00 2.50
COt1.PONENT
CONVERSION
----~---*'c::o .
-"'.ca- -- ----
1
2
III N HEXANE
1 BENZENE
80.59
. Oct ".." '."P
. ...0..0 "."0"0
---c:3_---"'-
-- -- -- ----
CATALYST VOLUME =
\ '. .
.TOTAL
.13173.1
90.0:)
FOq D2 CATALYSTS'" .THE FAC!; VF.LOC iTV IS.
25.6 FPM AT. 70 F.
CAT SPEC.
""J.J r'~.
. ''!U.,Df:N:
. .
VELOCITY
--------
-______C8
--------
D 2/
'0 ~
D r,.
'13
.~
1
25.5 .'
. 38.5.
51. 3'.
,g - \00
NO"; AT END
-------
ON AT 1~:D2:52 20,JUL 71 TUESDAY C5NRov
ASSISTANCE: 914-592-4R51 '
USER I Dj PASS':/ORD, p~OJF.CT I b--OALQ01, UOPROD,.
'145
SYSTEH--BASIC
V~R510N 14 JUN 71
. READY
LOAD tAT5
READY
RUN
12:39
CAT5
14:04
07/20171
ENTER CONVERSION,C,I'?; iNLET TF:t~~E!1ATURE, T1,' DEGRJ:ESF;
AVG GAS FLOW, G, SCFM.
190, 60J ,'10000
IF YOU II/ANT lHE SOLVEiH CODE PRH!TED, ENTF:R'1; IF NOT, 'O?O
.ENTER SOLVENT CODE NUMBERS. USE ZE~OS. TO 11AKE 8 ENT!"I,\::S
?14,1,0,0,0,0,0,0
'CNTER CO~JC~:~TRATIONS,
16.1,6.4~,O,0,O~0,O,O
y un ,
3TU/SCF.
USe: 1)'5
TO MAKE a ENTRiES
INLET TEMPERATURE, TI = 600
OUTLET TEMPERATURE, TO= 1225
DELTA T = 625
Co;~PONENT CONCENTRATIONS AT 5 POS IT I /)N5 l"J :lED FOl.LO"1
Y(l) Y(2) YO) Y( II) y(5) .V(I1) yeT) Y(~) x
-- -- - ----- ----- ----- ----- ----- -- -- - ----- -----
J=l 6.66 LI.17 .no .at) .oC) .00 .0f) .1')1) .~3
J=2 6.49 1. 86 .00 .00 .00 .O/) .00 '.00 .49
J=3 5.76 .22 .00 .00 .1)0 .00 .00 .00 .94
j=I! 3.68 .00 ~()O . a/) .;':)0" . . 00 ;01)0 .00 1.611
J-C:: 1. 30 .00 .00 .00 ~oo .00 ..0,), , .00 2.50
- .1
COt1PONENT
CONVE~SIO'J '
------.--.--
... -.---------
1
2
III N IiEXAf'!E
1 BENZEI-JE
go. 5'.> ,
:::::::: ::
----------
----------
CATALYST VbLUME =
TOTAL,
13J.7 31. 3
;I O. J9
FOR 02 CArALYSTS, TH,E FACE VELf1CITY is
';',7.3 Fpr-1AT 70,F
$-\0'
-------
--------
--...-----
--------
D 2
/) 3
D 4
.".
".".
?'f.3
:11. ;)
51: . r,
0,2
F1
N()'.-J AT END
111:09
RAN Q MINS
0.511 SECS
READY
RUN
CAT5
14:1:>
07/2rJ/71
ENTER CONVE~SION,C,%; INL~TT~HPERATUR~, Tl, DEGR~ES F;
AVG GAS FLOW, G, SCFM.
?90,f)')O,20000
I F YOU \oJANT TrlE S0LVENT CODE PR~NTEDp ENTER 1; I F NOT p 010
ENTER SOLVENT CODE NUMRERS. USE ZEROS TO MAK~ 8 ENTRIES
?14,1,O,~,O,O,0,0
ENTER CONCENTRATIONS, yeN), BTU/SCF.
16.7,6.42,0,0,0,0,0,0
USE O's Tn MAKE 8 ENTPT~S
INLET TEMPERATURE, TI = 600
OUTLET TEMPE~ATURE, TO= 1225
DELTA T = 525
cor~PONt::NT CONCENTP-ATI0N5 P.T 5 pns I Tt O~Jc; H! :3ED FnL U)\.f
y(1) y(2) y(3) yell) yes) yeG) YO) yeg) x
----- ----- ----- ----- ----- ----- ----- ----- -----
J=1 6.66 IL17 .00 .00 .0'1 .00 . ()o) .00 .21
J=2 6.119 1.86 .00 .00 . I)~) .00 .~):1 .00 .119
J=3 5.7r> .22 .00 .ao . ;JJ .00 .00 .00 .94
J=lJ 3.68 .00 , .00 .00 . ao .00 .0') . :1':J 1. ff
J=5 1. 30 .00 .00 .00 .')0 .00 .Or) ~oo 2.S0
COMP01\i~NT
COi'JVEr-S I f)N
----------
----------
1
2
14 N ,HEXANE
1 ~E"'Z ENE
81).59
~" ... ~. . ~ ...
".......,......,
----------
----------
CATALYST VOLUME =
TOTAL
263462.6
)0.19
FOR D2 CATALYSTS, THE, FACE VELOCTTY IS
27. 3 FPt-~ AT 70 F
CAT SPEC
"JUt1i1EP
VELnCITY
--------
--------
--------
D 2
D 3
D l~
'. ....
27.3
1:1 . ()
51: .6
", '"
--II'''
:: ,:
NO't! AT F.UD
II!: III
~AI'1 a :.11/J5
,) . 25 Sf. C 5
.!i-I 01..
-------
READY
OFF
OFF AT 1~:14 .
ELADSED PROCESSOR TIME.. 0 PlhS
ELAPSED TERMINAL TIM~... 12 MtNS
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8-103
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APPENDIX C
CALCULATIONS, QUOTES AND CORRESPONDENCE
ADSORPTION
TABLE OF CONTENTS
CONTROL PROCESS CONCEPT
ADSORPTION PROCESS CALCULATIONS
C-l
C-2
INSTALLED CAPITAL COST ESTIMATES
OPERATING COST ESTIMATES
C-23
C-38
ADSORPTION ISOTHERMS
VENDOR BUDGET QUOTATIONS
C-53
C-58
-------
PROCESS CONCEPT FOR CONTROL OF HEXANE AND BENZENE FUMES
FROM A PAINT BAKING OVEN VIA ADSORPTION
WITH SOLVENT RECOVERY OR INCINERATION
A paint baking oven exhausts a 50/50 weight percent
mixture of benzene and benzene vapors in air at a temperature
of 375°F and at concentrations of 15% and 25% of the lower ex-
plosive limit (LEL).
An economic evaluation of adsorption-recovery and
adsorption-incineration with and without heat recovery control
systems is performed for oven exhaust volumes of 1,000, 10,000
and 20,000 SCFM. Disposal (sewerage) costs of the small amount
(less than 5 gpm) of waste water were not included. These costs
would be small and would not affect the overall magnitude of
the estimated operating costs.
The system's basis of evaluation is at 90% recovery
or conversion of emissions. The oven is assumed to operate
2 shifts per day and 365 days per year. .
In addition, an economic evaluation of adsorption
and incineration without heat recovery is developed for an
oven exhaust solvent concentration of 200 ppm under the same
condftt6ns-out1i~ed above.
C-1
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CONSTRUCTION COST ESTIMATE
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July 26, 1971
Mr. Sam Martin
Singmaster & Breyer
235 East 42nd Street
New York, New York 10017
RECEIVED
AUG ...4 1971
Dear Sam:
Singmastcr 1& 6!reyer, Inc.
This letter will confirm our re~t conversation with regard to solvent
recovery of a hexane benezine stream at 100°F in atmospheric pressure.
, Enclosed are adsorbtion isotherms for hexane and benezine which should
be of valuable assistance to you in designing this unit as well as some other
information which will give you a good'.background on adsorption systems.
Normally on such a solvent system a workinE capacity of 10% is us~d for d~-
siEn purpJ?_~~!.u~h_I~"he.U..ey..e..,.y.OJ.L.wi 1,1 finn R highpr l'.
-------
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c-.s-t
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.!i:1'l8
~~~.
~.
ESTUIATED
BENZENE
ON BPL ACTIVATED CARBON
ADSORPTIOY OF
0.0001
0.001
ADSORPTION
PRESSURE.
0.01
0.1
PSIA
. .
c -.s-,r
"'M~
I((c!"
-------
July 29, 1971
Mr. Sam Martin
Singmaster & Beyer Co.
235 East 42nd Street
New York, New York 10017
RECEIVED
AUG - 2 1971
Sirogmaster 8: 8reyer, Inc.
Dear Mr. Martin,
As I mentioned to you in o~r telephone conversation of
this afternoon~ the equilibrium c~pacity of our Grade 256, 4 x.lO
for the benzene.-Q, hexane=air mixture-was 23% w/w. The conditions
studied were as follows: .
Concentration:
0.15% V/V benzene
0.15% V/V n, hex~ne
O. 70% V /V air
1000 + 20F
3" diameter x 3" height
Bed Temperature:
Bed Dimensions:
Flow Rate:
20
L/min.
Because the superficial velocity though the bed was low, and
because the bed was relatively shallow, wavefront data (breakthrough
and capacity at breakthrough) were not determined. However break-
throu h time was in excess of 6 hours and ca acit at this oint was
probably 18-20% w w.
The carbon was steamed for 15 minutes at 2750F (45 PSIA). The
bulk of the adsorbate was stripped after five minutes; this would
result in a steam solvent ratio of 4:1. As I mentioned, this ratio
should decrease markedly under plant conditions, where steam of
much higher pressure ~ 150 PSIG) could be used.
, c..S~
-------
-2-
Mr. Sam Martin
7/29/71
The relative concentrations of the benzene and hexane on
the carbon were determined by cold-trapping off the adsorbate.
The adsorbed phase was found to be hexane rich, the benzene
amounting to 30% of the desorbedmaterial, indicating preferential
adsorption of the hexane.
I hope these data are useful to you in designing your solvent
recovery system. If you have any further questions regarding Grade
256 or any of our other activated carbons, do not hesitate to call
me.
Sincerely.
C-57
-------
July 19, 1971
Singmaster & Breyer
235 E. 42nd ~treet
New York City, New York 10011
RECEIVED
JUl 21 1971
Attention:
Mr. Samuel Martin'
Sir:gmastleW 2& Brcy~rl' Inca
Gent1en:ten:
. .
We are pleased to confi~m the following approximate
prices for a 50/50 blend of hexane and benzene at 25% of the
L.E.L. at 375°F.
1000 cfm $35,000 + $10,000 installation
10,000 cfm $70,000 + $15,000 installation
20,000 cfm $110,000 -+ $30,000 installation
These prices are based on our standard compact design
which in sizes through 10,000 cfm are skid-mounted shop-assembled
units. Larger sizes must be field assembled.
We
information on
regret that we
of the various
enclose a copy of our brochures which give further
our various sizes and their capabilities. We
are unable to give you a detailed price breakdown
items comprising the system. .
Very truly yours, .
c:" .s" 8
-------
APPENDIX D
INCINERATION-ABSORPTION
AND
SCRUBBING-ABSORPTION
TABLE OF CONTENTS
PROCESS CONCEPT
INCINERATION-ABSORPTION
D-l
Process Calculations
D- 3
Installed Cost Estimates
DL 7
Operating Cost Estimates
D-10
SCRUBBING-ABSORPTION
,
Installed Cost Estimates
D-13
D-17
Operating Cost Estimates
-------
PROCESS CONCEPT FOR CONTROL OF HYDROCARBON EMISSIONS
BY INCINERATION-ABSORPTION SYSTEMS
AND SCRUBBING-ABSORPTION SYSTEMS
INCINERATION-ABSORPTION
The economic evaluation of incineration of waste
organics followed by absorption of soluble gases and par-
ticulates is based on the industrial case history described
in Appendix E. A waste liquid stream of kerosene contain-
ing DDT is incinerated by a direct-flame incinerator and the
incinerator gases are passed through a packed absorption
tower to remove soluble gases and particulates. In the
particular example studied, HCl was the principal waste gas
collected in the absorber.
Three levels of operation were considered: 100 gph
(2,000 scfm); 500 gph (11,800 scfm); and 1,000 gph (23,500
sCfm). Operating costs were based on a 2 shift/day operation,
330 days/year.
SCRUBBING-ABSORPTION
The econOmic evaluation of wet scrubbing absorption
of soluble organic waste gases was adapted from an actual in-
stallation for the control of waste gases from specialty or-
ganic chemical synthesis. Isopropanol vapors were used to
develop the. design basis for the absorber, although other
less volatile organics are known to exist in the. waste stream.
The basic economic study was made at the actual
plant capacity. Additional capacity levels were estimated
by employing the "six-tenths factor rule". Operating costs
were based on a 2 shift/day, 330 day/year operation.
Waste Water Disposal
Rather than attempt to estimate overage waste water
disposal costs (e.g., negotiated rate for sewerage), the op-
erating cost calculations considered a higher initial water
cost (process water instead of cooling water) and assumed no
disposal costs. For specific cases, the trade-off of inftial
water costs ~rid disposal costs should be treated on an in-
D-l
-------
dividual basis. In this generalized economic study, detailed
treatment of such trade-offs were not made because of the
strong dependency on local plant conditions. However, the
use of a higher initial water cost should approximately balance
off the cost of disposal.
0-2
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-------
APPENDIX E
INCINERATION-SCRUBBING SYSTEMS
FOR
HYDROCARBON EMISSION CONTROL
-------
ACKNOWLEDGMENTS
Appreciation is hereby expressed to the Maurice
A. Knight Company for their cooperation in permitting use
of their file data and design examples. In particular.
the support of Mr. M.A. Knightp Jr. p in allowing release
for publication of the pressure drop and f100ding corre-
lations. and the help of Mr. David Cooper in reviewing
th~calculations is particularly appret1~ted. Permission
of the Garver-Davis Company to use computer print~out data
for petrothemical waste incineration furnace design is also
duly acknowledged.
-------
I.
II.
I I I.
IV.
V.
TABLE OF CONTENTS
INTRODUCTION
FURNACE DESIGN
ABSORPTION TRAIN DESIGN
A.
Quench Tower Weight
1. Adiabatic Saturation
2. Cooling-Condensing
B .
Tower Area
1. Example Problem in Area Estimation
Packed Height Computation
1. Adiabatic Saturation
a. Height of Transfer
b. Number of Transfer
2. Cooling-Condensing
C.
Unit for Heat Transfer
Un i ts
D.
Pressure Drop Computation
1. Pressure Drop Estimation Example
DESIGN CASE HISTORY
CONCLUSIONS AND RECOMMENDATIONS
APPENDED TABLES
BIBLIOGRAPHY
Page No.
E-1
E-5
E-8
E-9
E-9
E-10
E-11
E-14
E-17
E-17
E-18
E-19
E-20
E-22
E-24
E-26
E-36
E-38-51
E-52
-------
Table
2-1
3-1
A-l
A-2
A-3
A-4
A-5
A-6
A-7
A-8
Fig u re
1 -1
2-1
3-1
LIST OF TABLES
Page No.
Incineration Combustion Products Analysis
Constants for Flooding Velocity Equation
E-6
E-13
3/4-lnch Ceramic Raschig Rings
l-Inch Ceramic Raschig Rings
E-38
E-40
1-1/2-lnch Ceramic Raschig Rings
2-lnch Ceramic Raschig Rings
E-42
E-44
3/4-lnch Ceramic Berl Saddles
l-Inch Ceramic Berl Saddles
E-46
E-48
1-1/2-lnch Ceramic Berl Saddles
2,-lnch Ceramic Berl Saddl,es
E-50
E-51
LI STOF FIGURES
Liquid Or~anicWaste Di~posal
Photo ,of Incinerator:Scrub~er. Installation.
of Section V Case-History . .'
. j
Page No.
E...3
Components of Packed Tower Pressur~ Drop
E-7
E-23
-------
SUMMARY
The design technology of incineration-scrubbing
systems for the disposal of organic and hydrocarbon wastes
under conditions avoiding the emission of secondary hydro-
carbon-derived contaminants has been reviewed and evaluated.
The scrubbing design techniques covered are those for siz-
ing equipment based on a sequence of adiabatic quench of the
hot combustion gases, followed by absorption of soluble gas
contaminants in a cooling-dehumidification section. The de-
sign approach to each of these sections is based on separate
and distinct heat transfer considerations. The absorption
of soluble pollutants is a secondary and incidental desi~n
consideration, because satisfying the heat transfer require-
ments for quench-cooling of the hot combustion gases gen-
erally provides for the necessary dissolution of soluble gas
contaminants such as partially-oxidized hydrocarbons and in-
organic acid gases.
A number of hitherto-proprietary precise industrial
design correlations and methods for the estimation of packed-
tower scrubber flow area and pressure drop are presented for
the first time. The use of t~ese techniques is demonstrated
for an industrial design case~history involving the inciner-
ative disposal of DOT-kerosene solution, with complete removal
of the HC1 formed on combustion. Because certain aspects of
the full design procedure are empirical and not rigorous,
information .has been provided on all necessary approximations
and safety factors required to achieve a feasible design.
This installation is in operation and meeting all design
specifications, and the techniques described are equally ap-
plicable to the design of scrubbers for the control of sol-
uble oxygen-containing hydrocarbons.
Although the area and pressure drop design methods
are highly accurate and reliable, considerable uncertainty
still exists regarding packed height estimation for cooling-
dehumidification operations. Because countercurrent contact
height is the primary factor controlling the degree of con-
taminant removal, and thus the level of emission, additional
research work is needed in this area, and general research
program recommendations are presented. .
-------
1.
INTRODUCTION
The purpose of this study was to identify and pre-
sent the technology involved in incineration-scrubbing systems,
with emphasis on the nature of the design problems and ade-
quacy of the data base involved in the scrubbing train-
portion of the system. Incineration of industrial organic
liquid sludges, and the combustion of municipal garbage and
other solid wastes is a popular method of disposal, and,
when due care is taken in system design to prevent air pol-
lution problems resulting from the combustion process, is an
acceptable disposal technique with regard to total emissions.
However, combustion of organic wastes, particularly with the
usual heterogenous feed compositions encountered, can result
in secondary organic and inorganic emissions which can be
serious health or material hazards. These include:
1.
Possibly carcinogenic polynuclear hydro-
carbons usually emitted in association
with carbonaceous particulates (Hange-
brauck, 1967).
2.
Acid gases resulting from halogen,
sulfur or nitrogen constituents of the
incinerated wastes.
3.
Partially-oxidized hydrocarbons such as
lIacids, anhydrides, 1actones and unsat-
urates due to thermal cracking and oxi-
dation in the combustorll (Sandomirsky,
1966).
Wet scrubbing of the combustion off-gases is fairly
effective for the removal of all three classes of possible
air pollutants listed, and both spray and packed tower
scrubbing have been employed for combustion gas treatment.
Because of certain inherent deficiencies of spray tower
scrubbing for multi-stage equilibrium contacting require-
ments, the packed tower is generally preferred for this
type of operation. In particular, packed tower technology
for the case of halogen-acid removal or recovery from a com-
bustion gas pre-dates air pollution control efforts, and
IIpackagell incinerator/packed scrubber systems have been avail-
E-1
-------
able for some years frOm both incinerator and packed tower
manufacturers. A schematic of this type of system is shown
in Figure 1-1, which is a double-tower unit required for
the complete removal of moderately-soluble acid gases such
as 502.
Whi~e the design of scr~bber systems is normally
predicated either on heat transfer or acid-gas dissolution
requirements, packed and spray scrubbers also have certain
capabilities for both particulate and oxidized hydrocarbon
removal. The use of irrigated packed towers for direct
particulate emission control has not been explored, although
it has been observed in a number of industrial installations
that the control level achievable is quite high. On the other
hand, the extent of removal of water-soluble oxidized organic
emissions in a scrubber system is directly calculable, and
the design methods reviewed in this report for acid-gas re-
moval apply equally to the oxidized hydrocarbon pollutants
in the combustion gases.
As indicated in Figure 1-1, the primary tower
serves as both a quench unit for the hot combustion gases,
and as a scrubber for the more readily-soluble gases. For
the case of chlorinated organics, or other materials which
will yield a highly-soluble gas contaminant on combustion,
the primary scrubber-cooling tower is all that is required
for complete contaminant control. This simplified single-
tower system is also highly applicable to contaminants com-
prised of partially-oxidized hydrocarbons because most of
this class of compounds have ready solubility in water.
Because the size of such installations is usually
small, and operation may be intermittent, the heat released
on combustion is not utilized for steam generation. However,
in several industrial installations, the waste sludge or
slurry to be incinerated contains sufficient water to war-
rant pre-concentration by using this stream as the primary
tower quench-scrub liquor prior to pumping it to the incin-
erator. This form of direct-contact heat utilization for
waste concentration effects a measure of fuel economy by
saving on the supplementary natural gas or other fuel re-
quired to sustain combustion, and it is expected that, where
applicable, this operation will see increasing use.
The secondary ~crubber:i~ 'generally used only for
difficultly-soluble gases w~ich require chemical reaction
for comp1~te removal. The most common case is the removal
of 502 with recycled dilute caustic solutions, where an in-
E-2
-------
COUR TESY OF
MAURICE A. K~IGHT
AKRON, OHIO
WAT[R
COMPi\NY
\
~
\-, (I;. '\>' it ~
.....it,..'. , ~ ..to ~ to.;
.... L I't' - "0>1'
6(~1:./~::'{",{',,~ l~ ..
''1. f,4 ~ "" ,. -,.-'"
~~\~ "Jti~3~'~
~~~"'f'?\f\ I~,p'~
~- Do L...."r'. .~~\..
\~~~)!ff..~ '1 ....~
4"..,\~.It'~+"'(' i'
~ t( ~",-:"!'('ti ~~
~.~,t:.~,~.J .:'~ ~
.":- ~~.\it:.' PI.!.. Z?
.':..;" '..~#~ ~..~..
,
"
WASTE
SURGE TANK,
SECONDARY SCRUBBER,
LIQUID ORGANIC
WASTE DISPOSAL
E-3
-------
itial batch of 8 to 10 wt.% NaOH is circulated until the con-
centration drops below 1 or 2%, at which point it is discarded.
Where secondary chemical reaction scrubbing is employed, the
exit gas temperature of the primary tower is controlled to
allow optimum mass transfer/reaction rates in the second tower,
and the entering temperature specification becomes a design
element on the primary scrubber. Alternatively, where the
exit gases from the primary scrubber pass directly to the stack,
steam plume or buoyancy considerations govern.
Because the problems of design of an old-in-the-art
pollution control device such as the packed scrubber provide
an insight into the general methodology and data base adequacy
of control technology, a detailed review of design procedures
was undertaken. Many of the design methods covered in this
review have not previously been published, but have been em-
ployed for industrial design problems for a number of years.
E~4
-------
I I .
FURNACE DESIGN CONSIDERATIONS
Furnace design is a highly-specialized and standard-
ized field~ and this aspect of the incineration/scrubbing
operation will not be considered in this report. However, the
temperature and composition of the combustion gases are the
primary working data for the design of the quench-scrubbing
system, and this determination is of major importance. The
industrial modus operandi is for the furnace manufacturer to
supply the scrubber manufacturer the detailed composition and
thermal information necessary for scrubber design calculations.
The separation of the areas of technical expertise between
furnace and scrubber manufacturers requires the preliminary
initiation of a high level of liaison and data exchange tn
order to arrive at the optimum overall design, and joint de-
sign responsibility is the "rul'.
Prior to the advent of the electronic computer,
the estimation of combustion gas temperature and compositions
was a tedious task 6" approximation. The computer, and the
availability of reliable chemical reaction equilibria data
from various rocket development programs, has simplified the
problem of composition determination, and has provided much
more precise information. A typical computer program is that
provided b¥ the Lewis Research Certter of NASA, available (with-
out charge) as the Lewis CEC, or Chemical Equilibrium Cal-
culation program. This is the program used by one furnace
manufacturer, Garver-Davis, Inc., and they have provided the
computer printout data of Table 2-1, taken from a typical in-
dustrial problem requiring the incineration of a waste.
chlorinated ,hydrocarbon stream. While compositional equilibria
are stated at several temperature levels, the true operating
temperature is a function of the firing rate, and this quan-
tity is specified by the furnace engineers. .
It should be noted that the furnace and primary
quench $crubber are most often built as an integral, com-
pact assembly. This is the result of the fact that ceramic'
materials are common to the furnace firebrick and the acid- 1
brick lining of the scrubber and the extended-surface pack-
ing it contains. Figure 2-1 is a photo of such an install-
ation with the furnace in the foreground and the scrubber
in the rear. This installation is the one for which the de-
tailed design example is presented below in Section IV.
E-S
-------
, r,ARVER-nAvtS, tNC. ,
POLLUnD',J CONTROL SYSTE"~
tNCINERATnR COIo4AtlSTION ?'OOUCTS ANALYSIS PREPARED FOR
~!'!~:"':"liTr":.r.JI&_~.
LAKE CHARLES. LA. '
SECOND STREA"'!
THERMO[\YNA~tC E~UILIE\qtllM PI'/OPr::~TI£" AT ASSIGNeD
TF.:Io4PERATURF:S MID PRE,?SIIJ;ES
"'ASTE
"'ASTE
WASTE
WASTE
WASTE
CHEMI CAL FOR'~IILA
C 2.00000 H 2.00000
C 2.00000 H 1.00000
C 2.00000 CL 6.00000
C 5.~0000 H 10.00000
C 3.75000 H 5.00000
CL 2.00000
CL,3.00000
CL 2.00000
CL. 1.c;\1000
~T FRACTION
I SEE flOT£)
.26042
.~1&67
.03125
.052n1\
.23~58
\.00000
OXIDANT 0
.21000
N
.70000
COMBUSTION CONDITIONS
FUEL TO WASTE wEIGHT RATIO:
.00000
OXIDANT TO WASTE wEIGHT RATIO=
PERCENT EXCESS AIR= 20.0000
4.07726
P. ATIo4
T. DEG F
1.000
1200
1.000
1400
1.000
1600
1.000
11100
1.000
2000
1o40LE FRACTIONS IN INCINERATOR COMBUSTION PRODUCTS
C02
CL
CLO
CL2
.141987
.000016
.000000
.020158
.111157/1
.000078
.000001
.01437/1
.1412118
.000268
.(001)02
.010:!l40
.141072
.000728
.000005
.007497
.140895
.001643
.OOOOOCJ
.n05444
HCL .082907 .094050 .1016l'2 .106711J .109752
H20 .033910 .028122 .024151 .0215111 .01CJ90b
NO .000005 .000017 .000050 .000117 .000240
N2 .695169 .693163 .691725 .690634 .6897011
OH .000000 .000000 .000000 .000002 .000001>
02 .025848 .028(,12 .0301193 .031708 .032396
ADDITIONAL PRODUCTS WHICH WERE COUSIDEPED BUT WHOSE "'IOLE FRACTI"NS WERE LESS THAN .0110005 FOR ALL ASSIGNED CONDITIONS
C CIS) CCL CCL2 CCL3 CCL4 CH CH2 CH3 CH4
CN CN2 CO COCL COCL2 C2 C2H C2H? C2H4 C2N
C2N2 C20 C3 CLCN CL02 CL?O H ,HtN ~CO H02
HZ HZOIL) H201S) N NH NH2 NH3 NOCL 'N02 H02CL
NZC N2H4 ~20 N204 0
NOTE. WEIGHT FRACTION OF WASTE IN TOTAL WASTE. OF FUEL IN TtlTAL FUEL. AND OF OVIDANT IN TOUL OXIDANT.
E-6
-------
,FIGURE 2-1
PHOTO OF INCINERATOR-SCRUBBER
INSTALLATION OF SECTION V CASE-HISTORY
E-7
-------
I I I.
ABSORPTION TRAIN DESIGN
The standard design approach to sizing the packed
tower absorption train on an organic waste incineration pro-
cess consists of four procedures:
1.
Weight Flow Calculation: The liquid flows
required for treating the given amount of
combustion gases are calculated from heat
and mass balances.
2.
Tower Area Estimation: Tower cross-sectional
flow areas are estimated from limiting flow
(flooding) equations and pressure drop con-
siderations for the calculated weight flow
rates of gas and liquid for several packing
sizes. The ultimate choice however is an
economic one.
3.
Packing and Tower Height Estimation: Packing
depths are calculated for the specified degree
of heat or contaminant removal from literature
or experimental data on transfer efficiency
of a unit height of packing, or volumetric'.
capacity coefficients, if and when available.
Pressure Drop Calculations: Pressure drop
through the absorption train is estimated
from av~ilable CBrrelations using the mass
liquid and gas flow rates and the result is
either matthed against, or specified for,
the heat output of the combustion air blower.
4.
The only uncpnventional part of the absorption-
train design procedure concerns the first tower. The
direct-contact quenching-cooling functions of the first
tower are primary, and this tower is therefore designed
on a heat tr~nsfer basis. In situations where a highly-
soluble acid-gas such as HCl is the major contaminant
specified for removal, a single-tower train is usually
sufficient, because enough cooling water can be fed to
the tower to serve both heat and mass transfer functions,
with only nominal flow area requirements. Thus, for
E-8
-------
chlorinated hydrocarbon disposal, quench-cooling tower design
is the complete system design, and the general procedures for
this unit are reviewed below.
A.
QUENCH TOWER WEIGHT FtOW CALCUlATION
The quench-cooling tower in an incineration/ab-
sorber train performs the following functions:
1.
Quenching the hot combustion gases to the
adiabatic saturation temperature by the
evaporation of water.
Cooling the gas from the adiabatic satur-
ation temperature to the desired stack
temperature.
2.
Completely absorbing the hig~ly-soluble
gas contaminants, and partially absorbing
the moderately-soluble contaminants.
These are independent functions, and the water flow require-
ments for each may be separately calculated.
3.
1.
Adiabatic Saturation
In order to estimate the quench water requirement
necessary for cooling the hot combustion gases to the adiabatic
saturation temperature, tas' the water content of the quenched
gases at this temperature must be determined. Because adia-
batic saturation is isenthalpic, the value of tas is obtain-
able from the enthalpy content of the combustion gases in
conjunction with an appropriate psychrometric table or chart.
The enthalpy of the combustion gas is given by the experi-
mental gross heating value of the waste, if available, or
from the temperature composition data of the equilibrium
furnace calculations (computer print-out). The enthalpy
value per pound of dry inert combustion gas is referred to
a standard air-water psychometric table, to obtain the water
content at saturation. It may be assumed that the difference
in molecular weight between the combustion gas and air will
have a negligible effect on the quench water rate so cal-
culated. The minimum evaporation water required, Wq' is
then calculated by:
Wq = Wg{Has - Hi)
( 3-1)
E-9
-------
where Wg = weight rate of dry gas flow, lb/hr.
Has = saturation humidity at tas' lb H20/
1 b dry gas
2.
Hi = initial combustion gas humidity,
lb H20/lb dry gas
Cooling-Condensing
If only the water of vaporization necessary to quench
the combustion gases isenthalpically to the adiabatic sat~
uration temperature were to be supplied to th~ quench tower,
the bottom section would obviously run dry, and the tower would
have no capability for sorption. It is therefore necessary to
supply water to the first tower in excess of the adiabatic
evaporation needs in order for it to function as a scrubber.
There ar~ several possible methods of calculating the total
water requirements of the quench tower, including:
( a)
(b)
( c)
Total Liquid Heat Load: The water rate is
determined by the maximum allowable exit
liquor temperature, under the total liquid
heat load and maximum inlet water temper-
ature. total liquid heat load is calculated
as the' sum of the gas enthalpy change over
the tower plus the heat of absorption and
dilution of dissolved solubles.
Minimum Sorption Liquid: The outlet liquid
temperature and HCl concentration is deter-
mined by trial-and-error for the two con-
straints of an inlet gas HCl concentration
greater than the equilibrium back-pressure
of HCl over the exit solution, and a total
enthalpy balance around the tower.
Steam Plume Considerations: Depending on
the climate and local conditions, it may be
necessary to limit the exit gas temperature
to minimize steam plume formation. If
ample~quantities of cold water are locally
available, the steam plume may be suppressed
by using the tower feed water to cool and
dehumidify the gas by direct-contact cool-
ing and condensation (Kalika, 1969).
E-10
-------
For the single-tower absorption case, the mass
transfer requirement of (b) is almost always satisfied by
the amount of' water needed for the secondary cooling fol-
lowing the quench as calculated from an enthalpy balance.
It is then only necessary to check that the partial pres-
sure of the soluble contaminant in the exit liquor, at exit
temperatures is less than the partial pressure of the con-
taminant in the entering combustion gas. This avoids tria1-
and-error mass transfer calculations, and also places the
primary emphasis on the thermal considerations, which con-
trol the design.
B.
TOWER AREA
The design value for the cross-sectional area of a
packed tower is estimated from the limiting capacity or
flooding velocity of the packing. Wh'i1e Igenera1ized" cor-
relations for flooding velocities are available, these are
stated in terms of arbitrary "packing factors" and their
use does not permit "tight" design, and frequently leads to
serious error (Prahl, 1970). In addition, the generalized
flooding and pressure drop correlations contain no correction
for such significant physical system parameters as liquid
surface tension, or greatly distort the effect of variables
such as liquid density. Because of these and other limit-
ations, alternate and more precise estimating methods have
necessarily been developed and employed. It should be noted
that rigorous design convention calls for setting the oper-
ating mass gas flow rate at 70% of the limiting flood gas
rate for the established liquid mass flow. Uncertainty in
estimating the limiting gas flow rate can be covered by
dropping the design basis to 40-50% of estimated flood, but
this is a poor design approach which results in an unneces-
sarily larger and more expensive tower. The 70% flood de-
sign will generally result in operation slightly in excess
of the load p6fnt, which is the ~ost efficient region of op-
eration with respect to mass transfer.
Use of the 70% flood limit as a basis for tower
design requires a precise mathematical statement of this
limit, and specific packing equations: theoretically based
and experime~ta11y-verified, were developed by Lerner (1951)
and later iridependen~ly confirmed by Howkins (1958). The
flooding equations are of the form:
(Go/~)F = k(l - mLoO.57)(a/aw)n
(3-2)
E-11
-------
where Go = gas mass flow rate, 1b/hr/ft2
Lo = liquid mass flow rate, 1b/hr/ft2
~ = 9as density correction
{Pg/0.075)O.5
Pg = gas density~ lb/CF
Ow = surface tension of water
factor m
k,m,n = constants
The constants, k and m~ are specific for each pack=
ing type, and size, and values of these constants are pre-
sented in Table 3~1. These constants have been evaluated
from the data of Lubin (1949), one of the few independent
studies of packed tower limiting flows. As indicated in
Table 3-1, for most packings there is an apparent change
in flow mechanism at a certain high critical liquid flow
rate, which is specific for the packing. Accordingly, it
is found that two equations, involving a low-range and high-
range set of constants, yield a better fit to the data than
one overall equation. The nature of this critical transition
liquid rate is examined in more detail in Section 3-4.
The estimation of tower area from the flooding
equations of Table 3-1 involves a limited amount of triA1-
and-error calculation. Although the initial value for
tower area for a given packing must be assumed, the use of
the 70% flood design base and the appropriate flooding
equation yields very rapid convergence on the correct de-
sign area for the second or third trial. The estimation
procedure based on the equations of Table 3-1 have therefore
proved amenable to programming for automatic calculation
for the simple desk-top electronic calculators. A magnetic
card program for area estimation based on the flood equations
of Table 3-1 is available from the Hewlett-Packard Company
Program Library for use on their 9100-B calculator. The
hand-calculations required without electronic-calculator
help for a single design are not extensive, and can be out-
lined as follows:
(a)
(b)
Choose a packing and assume a flow area, A1
From the given weight rate of liquid and
gas flows, calculate the mass flows, (Lo)l
and (Go)l, and Go/~) for the assumed A1~
E-12
-------
Packing Type
and Size
TABLE 3-1
CONSTANTS FOR FLOODING VELOCITY EQUATION
(Go/fD}F = k(l - m Lo 0057) (cr Icrw}n
Ceramic Raschig Rings
3/4"
1"
1-1/2"
2"
Ceramic Berl Saddles
3/4"
1"
1-1/2"
2"
n ,= 0058 for low range (LoIPL)
Ii = 1.0 for high range (LoIPL)
(Lol PL)
Range
2
mxlO
k
Low: <:. 13605 1685 3.53
High: :> 13605 1518 3028
Low: <::.. 144.5 1870 3065
High: "7 144.5 1420 2081
Low: <. 209 2118 2072
High: " 209 1500 1088
Low: ~ 240 2280 2033
High: '"? 240 2060 2.28
Low: <.. 160
High: > 160
1825 3.17
1405 2.70
2360 2064
2740 2013
Overall
Low: ~ 285
High: '"7 285
Low: <.. 300
High: /' 300
3110
2715
1.866
1.532
E-13
-------
( c)
Calculate (Go/~)F corresponding to (Lo)l
from the flood equation.
Calculate the ratio, R = (Go/~)l/[(Go/~)F]l
Assume A2 = (R/0.7)(Al) and repeat steps (b)
through le) until R = 0.7.
(d)
(e)
This procedure is perhaps best illustrated by an example.
1.
Example Problem in Area Estimation
In a certain paper~coating plant~ toluene is used
as the co&t1ng~resin solvent. Following the roll-coating
operation~ the solvent is removed by passage of the paper
web over heated rollsg and a hood serves to remove the tol-
uene and hot air. The exhaust gases from the blower serv-
ing the hood had previously been passed directly to the stack
at 130°F at a rate of 35,000 lb/hr of solvent-free air. It
is now proposed to prevent this emission by scrubbing in a
packed tower with a chilled (40°F) high molecular weight
hydrocarbon solvent with a specific gravity of 0.85 and a
surface tension of 32 dynes/em. The solvent-rich scrubbing
oil is to be stripped in a separate column with direct steam,
but the amount of surplus plant steam limits the oil circu-
lation rate to 1700 GPM. What diameter absorption col~mn
would be required for use with 1-1/2-inch ceramic Raschig
rings?
Trial 1: Assume Al = 100 ft2
L = (1700)(8.34)(b.85)(60) = 7,200 lb/hr/ft2
o (100)
(Lo/PL) = 135.8
From Table 3-1, (Lo/PL)c = 209, and low-range
equations apply.
(GO/~)F = 2118[1-0.0272(Lo/PL)0.57](cr/ow)n
~:
assume tower exit gas t = 60°F.
Average Gas Temperature = ( 130 ; 60)
PG = (29/359)(492/555) = 0.0716 lb/CF
= 95° F .
E-14
-------
T r1 a 1 2:
Trial 3:
4> = (0.0716/0.075)0.5'
4> = 0.978
(a/ow)" = (32/73)0.58
= 0.621
Now:
(Go)F = (0.978)(2118)[1 ~ 0.0272(135.8)0.57]
(0.621)
= (0.978)(2118)(0.553)(0.621)
= 711 1b/hr/ft2
(Go)act = 35,000/100 = 350 1b/hr/ft2
( Go) act / ( Go ) F = (350/'711) = 49. 3 %
A2 = (0.493/0.700)(100) = 70.4 ft2
Lo = (7,200/0.704) = 10,230 1b/hr/ft2
(Lo/PL) = (135.8/0.704) = 193
(GO)F = (0.978)(2118)[1 - 0.0272(193)0.57](0.621)
= (0.978)(2118)(0.436)(0.621)
= 561 1b/hr/ft2
(Go)act = 35,000/70.4 = 498 1b/hr/ft2
(Go)act/(Go)F = (498/561)(100) = 88.7%
A3 = (0.887/0.70)(70.4) = 89.2 ft2
(Lo/PL) = (193)(70.4/89.2) = 153.3
(GO)F = (0.978)(2118)[1 - 0.272(153.3)0.57](0.612)
= (0.978)(2118)(00537)(0.621)
= 691 1b/hr/ft2
(Go)act = (35,000/89.2) = 392 1b/hr/ft2
E-15
-------
Trial 4:
Tower Diameter:
(Go)act/(Go)F = (392/691)(100) = 56.8%
Take A4 = A2 ; A3 = 70.4 ; 89.2 = 79.8 ft2
(Lo/PL) = (135.8/0.798) = 170
(Go)F = (0.978)(2118)[1 - 0.0272 (170)0.57](0.621)
= (0.978)(2118)(0.491)(0.621)
= 632 lb/hr/ft2
(Go)act = (35,000/79.8) = 439 lb/hr/ft2
(Go)act/(Go)F = (439/632)(100) = 69.4%
Do = [(4/~)(79.8)]0.5
= (101.7)°.5
= 101-0"
E-16
-------
Choice of a proper packing and packing size is a
matter of some engineering judgment and experience, but the
above procedure will quickly reveal any flow-capacity inad-
equacy if the wrong choice is made. In actual practice, as
will be discussed below, the packing choice is also governed
by pressure drop limitations, and it is necessary to cross-
check the design mass flows for the head requirements once
the packed height requirement is known. It should be men-
tioned here that the calculator program containing the math-
ematical statement of the flood (and load) curves, also con-
tains mathematical equations giving the complete pressure-
drop ~. flow behavior for each of the packings of Table 3-1.
C.
PACKED HEIGHT COMPUTATION
The packed height requirement for the quench tower
is normally estimated as the sum of the two separate depths
corresponding to the two distinct and consecutive functions
of adiabatic saturation at the gas inlet section, and cool-
ing-condensing in the upper section. Alternate approaches
are possible, but these generally involve simplifying as-
sumptions as to the adiabatic saturation section requirements
and location. Despite pre-packing contact, it is good prac-
tice to assume quenching does not occur until the gas enters
the packing. Although the following treatment is not rigor-
ous, it comprises a conservative design procedure, utilizing
the best current information on transfer coefficients.
1.
Adiabatic Saturation
While the literature contains a number of studies
of heat transfer coefficients for the adiabatic saturation
of air with water, these have been determined only for a
very narrow choice of packings and low-temperature conditions.
Further, the absolute heat and mass transfer coefficient
values reported are not always corrected for packing end-
effects, which are relatively large for the shallow packed
depths necessarily employed to avoid equilibrium saturation
under laboratory conditions. For example, Hensel and Treybal
(1952) reported extensive data for 1-1/2" Berl saddles, and
when the data were extrapolated to infinite tower height,
the heat transfer coefficients underwent a three-fold re-
duction from the observed shallow-bed.va1ues. Thus, certain
manufacturer's test data on shallow beds should not be used
without correction for the re1ativel~ large and variable
end-effects.' There also appears to be a high degree of dis-
parity between field experience on gas-liquid contact time
and area requirements for the equilibrium quenching of hot
combustion gases, and the highly optimistic laboratory re-
E-17
-------
su1ts on low-temperature transfer coefficients given in the
literature. For example, for spray chamber quench units
employing recirculated water, design practice (Ka1ika, 1969)
calls for only 60% evaporation of the liquid injected into
the gas. For packed towers it has been found that a heat-
transfer (HTU)(NTU) approach yields comparatively conserva-
tive contact depth estimates and by conversion, allows the
use of the extensive literature (HTU) values for mass transfer.
( a )
Height of Transfer Unit for Heat Transfer:
The correlation that allows conversion of the lit-
erature values for mass transfer, (HG)d' to the heat transfer
values of (HG)h, is that of WilRe et al (1963), which is
stated for the case of gas-film resistance controlling as:
(HG)h = (HG)d(D/a)0.5(~d/~h)0.27(Ph/Pd)0.84(ah/ad)0.925 - 0.262
(3-3)
Where:
DG = mass diffusivity, sq ft/hr
a
~ thermal diffusivity, sq ft/hr
= mass liquid flow rate, 1b/hr/ft2
L
p
= gas density
= liquid surface tension
a
~
= gas viscosity
h,d = subscripts denoting heat and mass
transfer, respectively.
The group, ~G/a) in Equation (3-2) is somewhat dif-
ficult to evaluate as written, because direct data on thermal
diffusivities are not readily available. However, by defin-
ition: a = (k/pCp) (3-4)
Where: k : thermal conductivity
Cp : heat capacity
and the group, (DG/a) = (Cp~/k)/(~/pDG) = (Pr)/(Sc)
Where: Pro: Prandtl Number, (Cp~/k)
Sc : Schmidt Number, (~/pDG)
(3~5)
E-18
109(L)
-------
Because the moles of water/mole dry gas at the adiabatic
saturation temperature, ta$~:c6ndit;ons prevailing in the
gas film adjacent to the '1quid is generally greater than
2/1, the Prandtl number for steam is more truly descriptive
of thermal transport than that for air.
The value of (HG)d in the Equation (3-2) may be
evaluated from the data of Fellinger (Sherwood and Pigford,
1952) for the ammonia-air~water system, and converted to
the water-air-water system by the relationship:
(HG)H20 = (HG)NH3 (SCH20/scNH3)0.5
(3-6)
or
(HG)H20 = (HG)NH3(DNH3/DH20)0.5 (3-7)
Inasmuch as the diffusivities and Schmidt numbers
for water and ammonia are virtually identical, with SCH,O =
0.60, and SCNH = 0.61, the (HG)NH values for the reference
NH3-air-water ~ystem can be used dfrectly for (HG)H20 with
less than 2% error.
(b)
Number of Transfer Units:
The number of transfer units for heat transfer
based on gas-film resistance controlling (NG)h' may be
stated as:
t2 - tl
(NG)h = (l1t)m
= 1 haZ
CpG
(3-8)
(3-9)
where tl' t2 = 1 eavi ng and enter; ng gas temperatures,
respectively
(l1t)m
ha
= 10g mean temperature driving force
= gas film volumetric heat transfer
coefficient, Btu/hr/ft3/oF
Z
G
= packed height, ft
= mass gas flow, lb/hr/ft2
E-19
-------
In order to assess the relative magnitude of the adiabatic
cooling NTU requirements, if it is assumed that a hot com-
bustion gas must be cooled from 1800°F, to an adiabatic sat-
uration temperature of 180°F, using water at 140°F (assumed
constant) the transfer unit requirements are then calculated
to be:
(NG)h = In(1660/40) = 3.73
For stagewise contacting, assuming that (approximately)NTU c
NTP, the requirement for close to four equilibrium stages in-
dicates that, despite what may appear to be a large driving
force, the multi-stage requirements are significant. For the
case of the spray chamber, relative gas-liquid velocities
vary and are unstable, and departure from true countercurrent
flow can be caused by gas reciruclation due to liquid jet
momentum. Thus, for the example shown, if the true evapor-
ation efficiency value of Kalika (1969) mentioned previously
is used, then a total of seven spray stages would be required
to satisfy the indicated NTU.On the other hand, for a
packed tower, with true differential countercurrent contact,
assuming a representative order-of-magnitude value for (HG)h
of 1 foot, the same transfer can be effected in less than four
feet of packed depth. Thus, as has been shown by Pigford and
Pyle (1951), where conditions allow the use of the packed
tower, this is a more effective contacting unit than a spray
chamber.
2 .
Cooling-Condensing
Following passage through the evaporative cooling
section at the bottom of the tower, the gas then flows through
the upper section where, if the feed water is below the
adiabatic saturation temperature, the gas is cooled and part
of the water evaporated in the bottom section is condensed.
Unlike the adiabatic evaporation process, which is unequivocally
gas~film controlled, the cooling-condensing section involves
liquid-phase resistance to heat transfer. The degree of this
additi~nal resistance, and its effect on design assumptions,
has b~en the subject of a great deal of dispute but little
clarification. Further, the chances for fog formation in
this section are extremely good, and there is no method for
handling calculations under these conditions.
Packed height estimation for the cooling-condensing
region is a highly uncertain procedure, primarily because of
the lack of performance data. The reasons for this lack were
specifically noted and discussed by Treybal (1955) but no
E-20
-------
remedial action has been taken in the 16-year interim. It is
known that a certain amount of proprietary performance inform-
ation is on file and is utilized for design purposes by tower
fabricating companies, but the reliability of this data is
not known. The available published heat and mass transfer
coefficients give inconsistent results, with height values
varying from impractically-shallow beds of less than 6 inches
depth, to heights 'in excess of 20 feet for a given problem,
depending on which coefficient data and procedure is used.
It would be highly desirable to have a set of HTU values for
several packings, determined simultaneously on the basis of
enthalpy-potential, temperature, and humidity driving forces,
for the dehumidification of air by cold water. In the absence
of such data, the method that has been found to give the most
rational results for height estimates in this cooling-dehumid-
ification region is to assume that the liquid film resistance
controls. On this assumption, the NTU is given by:
(t2 - tl)
(NL)h = (T2 - t2) - (Tl - tl)
(t2 - t2) (3-10)
(Tl tl)
where
T = gas temperature
t = water temperature
1,2 = subscripts denoting top,and bottom of packed
section, respectively
As was the case for the gas-film coefficients, the
liquid-film coefficients are obtained by use of the equations
of Wilke, et al (1963) for converting (HL)d literature data
on the desorption of oxy~en from water to the corresponding
heat transfer value, (HL)h:
(HL)h = (HL)d(D/a)0.5(~h/~d)0.55(Pd/Ph)0.329(crd/crh)0.554-0.157
( 3- 1 1 )
10g(L)
Because of the fairly large uncertainties in the use
of liquid-film coefficients in the absence of direct support-
ing data, it is necessary to use adequate safety factors. It
is usual practice to use a 1.5 multiple (50% safety factor) of
the total theoretical packed height of both quenching and de-
humidification sections to cover the inadequacies in the basic
coefficient data.
E-21
-------
D.
PRESSURE DROP COMPUTATION
Estimation of the flow resistance of the scrubbing
train is necessary in order to specify the heat requirements
for the combustion air blower on the furnace. As shown in
Figure 3-1, the pressure drop through a packed tower consists
of several losses in addition to the "intrinsic" pressure
drop through the packing itself. However, while these sec-'
ondary head losses cannot be neglected, for packing depths
of more than a few feet, the intrinstC pressure loss of the
packing itself is the major item in the head requirement in-
ventory.
Although partial correlation of packed tower pres~
sure drop was achieved a number of years ago (Treyba1, 1955)
these early efforts failed to cover the higher liquid rate
range where the slopes of the pre-load log-log plots of 6P
vs. mass gas velocity i~creased above the normal value of
1.8. The M.A. Knight Company has accomplished a complete
correlation of the full pressure drop behavior of packings
on the basis of defining this latter critical liquid velocity
(termed the "gas-sweep" velocity) as a demarcation line wh~ch
horizontally divides the 6P vs. Go family of curves into two
zones of operation: pre~critica1 and super-critical. The
flooding line on the log-log 6P plot sets the upper limit to
the curves, while the loading line, which can be described
mathematically in a manner simi1ar to the flooding equations
of Table 3-1, serves to vertically delineate the graphs into
pre-loading and load-to-flood zones of operation. There are
thus four distinct zones of pressure drop dependency on flow
rates in a packed tower, and a mathematical description of
the pressure drop behavior specific to each zone is derivable
from analysis of the two-phase flow mechanics characteristic
of each flow zone. The flow zones and generalized specific
6P expressions are:
1.
Pre-Load, Pre~Critical
(6P/Z) = a1(Go/c/»1.8
2 .
Pre-Load, Super-Critical
(6P/Z) = a2(Go/c/»1.8 + Kl
Load-to-Flood, Pre-Critical
3.
(6P/Z) = a3(Go/c/»3.0
E-22
-------
---- PACK I NG
SUPPORT,
GRILLAGE
LIQUID
...- - .'.0,,_-
OUT
0A$
OUT
r
6PC... )
Pfg out
~Pdistributor
't--
.6Pexit
V-WEIR
~PpaCking
(INTRINSIC)
~p entry
r'
~P(PiPing)in
L
.h4!!
GAS
IN
FIGURE 3-1
COMPONENTS OF PACKED, TOWER PRESSURE DROP
E-23
-------
Load-to-Floodt Super-Critical
(AP/Z) = a4(Go/~)3.0 + K2
The fundamental utility and elegance of this cor-
relation approach is that all of the zonal constantSt a and
K are functions of liquid mass flow rates and packing type
and size, and the relationships may be theoretically explained.
The intrinsic AP equations for Raschig rings and Berl saddles
are presented in the Appendix to this report. The use of
these equations is best illustrated by an example.
4.
1.
Pressure Drop Estimation Example
For the example tower area design problem given in
Section B-1, the mass flows for 1-1/2" Raschig rin9s in a
10'-0" diameter tower are:
Go = (439)(79.8/78.5) = 446 lb/hr/ft2
(Lo/pL).= (170)(79.8/78.5) = 172.8
(Lo/pL)c = 209t Lo is pre-critical
From Table A-3, Appendix:
Loading:
(Go/~)L ~ 1070[1-2.14 x 10-2(Lo/PL)0.62]
(Go}L = (0.978)(lQ70)(0.477)
= 499 lb/hr/ft2
(Go)act/(Go)c =
so (Go)act
From Table A-3:
(466/499)(100) = 89~4% of load
is pre-load
Zone 1: Pre-Load, Pre-Critical Equation applies:
(AP/Z) = al(Go/~)1.8
where al = 1.45 x 10-6(102.60 x 10-3(Lo/PL)
and for (Lo/PL) = 172.8,
al = (1.45 x 10-6)(2.83)
E-24
-------
(l~P/Z)
= 4.10 X 10-6
= (4.10 x 10-6)(6.11 x 104)
= 0.25" H20/ft of packing
E-25
-------
I V ,
DESIGN CASE HISTORY
CASE HISTORY NO.1
The client's disposal problem involved the incin-
eration of 100 gallons/hour of 5% by weight DDT solution in
kerosene. The gross heating value of the solution was ex-
perimentally determined as 20,700 Btu/1b~ and the density of
the mixture as 50 lb/ft3. Removal of the HCl formed on com-
bustion is desired, and cold, fresh water is readily avail-
able at the plant site, with a maximum summer temperature of
50°F. The allowable maximum waste water temperature is 175°F.
1.
Material Balance
Bas is:
1 Hour
Total Weight litllJiCLBurried ';: (1..00 gal/hr)j50 lb/CF)
(7.48 gal/CF)
= 668 lb/hr
C1
Composition wt % lb/hr M. wt Mole/hr ]b atom/hr
DDT: C14HgC15 5 33.4 354.5 0.0944 0.04720
Kerosene: CO.8SHO.15 95 634.6 10.35 61. 3
668.0 61.3944
2.
Flue Gas Analysis and Flow Rates
The combustion gas analysis was obtained directly
from the computer program on thermodynamic equilibrium for
the furnace combustion products (Lewis, NASA CEC Program).
For this case, at a temperature of 1600°F, the following
flue gas composition was calculated. '
E-26
-------
Component Mole Mole/hr. . lb./hr.
%
C02.. .. 16.5948 54~2 2385:
N2 78.2904 254.5 7120
02 3.4679 . 11,,3 362
NO O. 0l3r 000427 102
. OH- 0.0001
C12 0.0012
(Inert) 98.3675 320.678 9868.2
HCl 0.1448 O. 4720 17.22
Total (Dry)
98.5123
321015 (BDG)*
9885.42
H20
1.4874
4.85
87.3
Total (Wet)
99.9997
326.0
9972.7
*Bone Dry Gas
3. Adiabatic Saturation Temperature
The value of tas is determined as that corresponding to the gas enthalpy content.
Heat Release
= (20,700 Btu/lbo)(668 Ib./hr.)
= 13, 800, 000 Btu/hr.
Enthalpy Content = 13,800,000 Btu/hr.
9885.42 Ib./hr. BDG
= 1397 Btu/lb. BDG
E-27
-------
From psychrometric table in Perry (1963) p. .!~:7, by interpolation:
tas = adiabatic saturation temperature = 19l.2oF..
water content at tas = 1.184 lb. li20/lbe BQG
saturated volume at tas = 47..27 ft.3/lbe BDG
4. Adiabatic Saturation Water Requirement
Ibs. Water EvapQX'ated/hr..
= (l.184)(9868q2) ~ 87..3
= 11,690 lb./hr.
5. Feed Water Rate (Total)
Based on steam-plume considerations (see Kalika (1969) and Crocker (1968»
and local meteorological conditions at the plant site, the exit saturated gas temperature
from the tower will be limited to 130oF. maximum.
At 1300F ., Saturated
Gas Enthalpy
= 155.9 Btu/lb. BOO
Gas Volume
. = 0.116 lb. H20/lb. BOO
= 17.516 CF /lb. BOO
Water Content
Overall Net Tower Heat Load
(a) Dry Gas
= 1311 800, 000 - (155.9)(9868)
= 12,262, 000 Btu/hr.
(b) HCl: Heat generated by absorption (Perry, 1963, p. k 137)
Ht.. Formation HCl (aq.)
= -39.85 kcal./g. -mole
Ht. Formation HCl (g.)
= -22.063
Ht. Soln. HCl
= -17.787 kcal./g. -mole
E-28
-------
Ht. Soln. HCl/lb.
= ( -17. 787)( 1800)
36.5
= 877 Btu/lbo
Soln. Ht.. Generated
= (877)(17022)
= 150 080 Btu/hr 0
Total Heat Load
= 12, 2621) 000 + 150 080
= 12p 277, 080 Btu/hr 0
Water
t = 175° -SOop" = 12SoF 0
Feed Water
= 12, 277, 080 Btu/hr.
(125)
= 98,250 lb. /hr.
= 196.5 GPM
Water leaving Adiabatic
Saturation Section
= 98,250 - 11,690
= 86, 560 lb. /hr.
= 173,,1 GPM
6. HCl Absorption Check
Assume H20 enters and leaves adiabatic quench section at 175°F.
Therefore, equilibrium check on HCl will be at 175°F. liquid vs. 191°F" saturated
combustion gas.
PHCl in gas = moles HC1/hr 0
moles BDG/hr. + moles H20/hr.
=
(0.472)(760)
(321.15 + (11, 777.3/18»
E-29
-------
PHC1
= 0.368 mm. Hg
Liquid
Concn.
=., 17.22 (100)
. 98, 250
= 0,,01752 m. %
From Perry (1963) p. 14-69 @ 80oe../) 2% by weight HCl has an equilibrium
P*HCl = 0,,0245 mm. Hg
(PlfCl) gas > prllHCl D ~nd absorption driving force is positive
throughout tower.
7.
Tower Area
Design is for 7rP/o of flood, using general equation
(Go/'/)p = k (1 - m(Lo/fL)0.57)(<1" /0' w)n
Because the lower the '/) - value, the lower the Go required for flood, design
is based on the lowest tower'/) - value, which corresponds to gas density at the top end
of the adiabatic saturation section, which is also the point of maximum weight rate of
gas flow.
gas densityp sat'd. @ 191.2oP. = (2.1841bs.)
47.27 CP
= 0.0462 Ib./CP .
'/)
= (O,,046~ 0.5 = 0.785
0.075
E-30
-------
Assume 1-1/2" Berl Saddles: (low range (Lo/f>L»
(Go to larger size if final A P is too high,,)
From Table 3-1
k = 2740
m = 2..13 x 10-2
and
for aqueous system (0"'/(1' w) = 1,,0
(Go)p = (0,,785)(2740)(1 - Oo0213(Lo/!>L)0057)
max", liquid rate
= 98, 250 1bo /hr.
Trial No.1: Take Area = 30 fte2
Lo
= 98,250 = 3,2751be/hr./ft.2
30
(GO)F = (0.785)(2740)(0.796)
= 17131bo/hr./ft.2
and
(GO>act = 9885 + 11, 777
30
= 7221b./hr./ft.2
R = «Go> act/(GO> p)(722/1713)
= 00422
too low.
Trial No.2: Take A2
= (0.422/0.7)(30)
= 18..1ft02
Adjust A2 = 19063 fto2 for 5' Diam" Tower
(Lo/fL>
(Go)F
= (98,250) = 5000 lb./hr./fto 2
19..63
= (5000/62.4) = 8001
= (00785)(2746)(0074)
Lo
= 15921bo/hro/ft02
E- 31
-------
(Gohct
::: (722)(30/19.63)
= 1l041b./hr./ft" 2
R ::: (Go)act/(Go) F = (1104/1592)
::: 0.694
Diam. :::
5' -0" for 1-1/2" Bed Saddles
(Note~ On review of designo client requested 50% excess combustion gas
scrubbing capacity to provide for possible peak o:!Lganic disposal
rate, and accordingly, diameter was revised to 6' -0" .)
For 6'-0" Diam., A ::: 28.3 ft.2
Lo
=
(5000) (19863)
28..3
:::
34701b./hr./ft.2
Go
:::
(1104) (1906~
28.3
=
765 lb. /hr . /ft. 2
8. Packed Height
A.
Adiabatic Saturation Section, Gas-Film Controls
Assume gas enters and leaves section at 1750F.
(NG) h = t2 - tI ::: In (1425)
(At)m . 16.2
::: 4.47'T:ransfer Units
(HQ>h: for: Lo ::: 3470e> Go ::: 765" from Fellinger data, Sherwood
and Pigfo;rd (1952)..
(HQ>h
= 0.9 ft.
::: (0.90)(0.60/0.61)0.5 = 0.90 ft.
::: (0. 9O)(D/~)0.5(}ldl )lWO.27(fh/ fJ&0.84(~h/a' &0.95-0.262 log Lo
(Ha>NH3
(HG)H20
E-32
-------
Temperatures
Parameter 70oF. 1300F 0
U air, cp 0.018 0.. 0192
o . (wet) 0.0742 0.0636
air
0H20 72.8
1750F"
1910F 0
0..0205
0<)0517
0.. 0462
63"
Because 0<>262 log Lo > 0.,9258 0 correction drops out
and
(HG)h = (0..90)(1.291)( 0,,018 ) 0..27{0.,0462) 0084
0.0205 0.0742
= (0.90)(1«>291)(0.965)(0.67)
:;:t 0.75 ft.
Zas
= (HG)h (NG)h
= (0.75)(4.47)
Zas
= 3.35 ft.
B. Cooling-Condensing
Design convention for the case where the liquid A t ~ gas At through
the cooling section assumes that liquid film controls. In this case:
(N L)h
= (t2 - tI) , In (T2:'~ t2)
(T2 - t2) - (Tl - tl) T1 - tl
(175 - 50) 80
= (130 - 50) - (191.2 -175) In (16.2)
= (125 ) (1 6)
63.8 .
= 3.13 Transfer Units
(H L)h Estimation
from the Wilke-Cheng (1963) correlation:
E-33
-------
(HL)h = (HL)d (D/oC) ° 0 S(Ph/ P&o.ss(f>dI f>W.329(°o5 = (100/0060)°,,5 = 10291
(Ph/ )1<1>0055 = (0..018/0..0185)°055 :: 00986
(~d/ f>h)0.329 = (0.0742/0.0668)°0329 = 1.035
(HL)h = (0.95)(1.291)(0.986)(1.035)
= 1.25 ft..
Packed Height = (1.25)(3.13)
ZL = 3.92 ft.
Total Theoretical Packed Height = (Z)as + (Z)c
ZT = 3.35 + 3.92
ZT = 71[)27 Fto
Using a 50% Safety Factor: Packed Height = 11.5 Ftc
9. Pressure Drop
~P for packing
Lo/fL = (3470/6204) :: 55.7
Operation with respect to loading
(Go/~)L = 1454 (1 - 1.. 755 x 10-2 (55.. 7)0.55)
= 1454 (0084)
= 1222 Ib./hr./ft.2
E-34
-------
so
(765/1222) (100) = 62.5% of loading Go
(/).P/Z) = al (Go/~)1.8
and from Appendix:
al = 7..90 x 10-7 (102..13 x 10-3(Lo/fL»
al = 7.90 x 10-7 (100.,18~
= 1,,202 x 10-6
(6P/Z) = (1..202 x 1O-~ (765)1.,819 fD = 100
= (1.202) (0.156)
= 0.1875 in. H20/Ft.
for entrainment separator section:
(6P/Z)dry = 4.28 x 10-7 (Go/~)1.8
= 0.0668"
(a) 6 P Packing, intrinsic = (0.1875)(11.5)
(b) 6,P Entrain. Sep"p 9" of 1-1/2" B.S. = (0.75)(0.0668)
*(c) AP Packing Supports, for (a) & (b) = (2)(0.11)
*(d) .6 P Distributor
(e) Entry & Exit Loses
Allowance for possible fouling = 50%
and (1.5)(2.94)
E-35
= 2.16"
= 0.05"
= 0.22"
= 0.11"
= 0.40"
2.94" H20
= 4.40" H20
Total
-------
v.
CONCLUSIONS AND RECOMMENDATIONS
Review of the incinerator-scrubbing design tech-
nology, as applied to the combustion and possible emission
of hydrocarbons, indicates that the data design base is very
unevenly developed. New and highly accurate design cor-
relations are available for estimating the packed-tower dia-
meter and packing pressure drop, while on the other hand,
estimation techniques for countercurrent contact path or
packing height are grossly inadequate because of the lack
of a design data base.
Tower design for waste combustion systems consists
of two distinct heat transfer problems: sizing and adiabatic
quench section and sizing the subsequent cooling-condensing-
absorption section. Both sections normally are encompassed
by a single scrubbing-quench unit, but the design approach to
each section is different. The design methods explored in-
clude estimation of the height requirements for the adiabatic
section using an NTU based on temperature driving force, to-
gether with the assumption that gas-film controls. Contact
height estimation for the direct-contact cooling-condensing
region of a scrubber is made on the assumption that liquid-
film resistance to heat transfer controls, an assumption that
empirically yields rational contact heights, but which is
theoretically unjustifiable {Wilke et a1, 1963}.
Despite the widespread use of direct-contact coo1-
ing-condensing-absorbing equipment for the control of emis-
sions from incinerator off-gases, the data base available
for rational design is virtually non-existent. Remedial re-
search programs are needed in the following areas:
1.
Experimental determination of HTU for
direct-contact countercurrent gas-water
contacting equipment, with simultaneous
determination of the HTU for the follow-
ing three different transfer potentials:
{a}
~ ~ ~
Enthalpy
Humidity {partial pressure of H20}
Temperature
E- 36
-------
2.
Inter-correlation of overall mass and
heat transfer coefficients on the basis
of individual film resistancesg so that
direct-contact overall coefficients can
be estimated from presently=available
literature data.
3.
Determination of the effect of liquid-
film heat transfer resistance in direct-
contact cooling-condensing operations
over appreciable temperature ranges.
4.
Experimental determination ~f the effec~s
of fog formation on heat and mass transfer
coefficients in direct-contact cooling-con-
densing operations, and on the efficiency
of particulate removal in such operations.
E-37
-------
TABLE A-I..
..
INTRINSIC PRESSURE DROP EQUATIONS
3/4- INCH CERAMIC RASCHIG RINGS
Critical Gas Sweep Liquid Rate: (LpIPL) = 136.5
. .
Loading: . (Go/~)L t = 674(1- 1.493 x 1O-2(Lo/pi.,)0.62)
Flooding:
Low Range: "
(L olfL> <. 136.5
'(Gol!l»F = 1685{1 - 3..53 x 1O-2(Lo/fL)0.S7)(cr / 136.5,
. .
(Go/f/)F = 1518(1 - 3.28 x 1O-2(Lo/PL)0.57)(0'" /0'" w)n
n = 0.58 for (LonoL) <.136.5
n = 1.0 for (LoIPL) > 136.5
Zone 1.
Pre-Load and Pre-Critical:
where
(Go/~) < (Go/~)L' ; (LoQOL) <. 136.5
(AP/N) = al (Go/~)1.8
al = 2.64 x 10-6 (104.83 x 1O-3(LolfL»
Zone 2.
Pre-Load and Super-Critical:
where
(Go/~) < (Go/~)L' ; (LolfL) > 136.5
(AP/N) = a2 (GolfD)leS + Kl
a2 = 2.18 x 10-6 (105.12 x 1O-3(Lo#>L»
K1 IIC 5.15 x 10-6 (Lo/PL)I.90
E-38
-------
3/4- INCH CERAMIC RASCHIG RINGS Cont'd.
Zone 3. Load-to- Flood and Pre-Critical:
where
(Go/'/)F > (Go/f/) > (Go/'/)L I ; (Lo/PL) <: 136.5
(AP/N) = a3 (Go/'/)3.0
a3 = 1.01 x 10-9(106.47 x 1O-3(Lo/'pL»
Zone 4. Load-to- Flood and Super-Critical:
where
. (GO/'/)F > (Go/'/) > (GO/'/)L' ; (Lo/PL) > 136.5
(AP/N) . = a4 (Go/'/)3.0 + K2
a4 = 8.64 x 10-10(106.38 x,lO-3(LoU>L»
K2 = 3.68 x 10-3 (LofL)0.75
E-39
-------
TABLE A-2
INTRINSIC PRESSURE DROP EQUATIONS
1- INCH CERAMIC RASCHIG RINGS
Critical Gas Sweep Liquid Rate: (Lo/fL>C = 144.5
Loading: (Go/f)L = 935(1 - 2.44 x 1O-2(Lo/t'L)0,,621
High Range:
(Lo/fL) <: 144.5
(Go/'/)p = 1870(1 - 3.65 x 1O-2(Lo,fL)0.57)(a" /cr'w)n
(Lo/fL) > 144.5
(Go/'/)F = 1420(1 - 2.81 x 1O-2(Lo/fL)0.57)(c:r/a"w)n
Flooding:
Low Range:
n = 0.58 for (LoqaL) L. 144.5
n = 1.0 for (LoIfL) > 144.5
Zone 1.
Pre-Load and Pre-Critical:
where
(Go/'/) < (Go/'/) ; (LoJOL).( 144.5
&..;
(AP/N) = a1 (Go/'/)1.8
a1 = 1.29 x 10-6 (106.60 x 1O-3(Lo/PL»
Zone 2.
Pre-Load and Super-Critical:
where
(Go/'/) < (GO/'/)L; (Lo/PL) > 144.5
(ta.P/N) = a2 (Go/f/)1.8 + K1
a2 = 2.57 x 10-6 (103.67 x 1O-3(LoJOL»
K1 = 1.70 x 10-6 (LoUOL)2.24
E-40
-------
1- INCH CERAMIC RASCHIG RINGS Cont'd.
Zone 3. Load-to-Flood and Pre-Critical:
where
(Go/'/)F > (Go/'/) > (Go/'/)L; (Lo/fL) <. 144.5
(AP/N) = a2 (Go/'/)3.0
a2 = 6.95 x 10-10 (106.42 x 1O-3(LoQ>L»
Zone 4. . Load-to- Flood and Super-Critical:
where
(GO/'/)F > (Go/'/) > (Go/'/) ; (LoIPL) > 144.5
(AP/N) = a3 (Go/'/)3.0 + K2
a3 = 9.11 x 10-10 (104.77 x 1O-3(Lo/,oL»
K2 = 3.63 x 10-4 (Lo/fL)1.35
E-41
-------
TABLE A-3
INTRINSIC PRESSURE DROP EQUATIONS
1-1/2-INCH CERAMIC RASCHIG RINGS
(3116" WALL) .
Critical Gas Sweep Liquid Rate: (LolfL) ~ 209
Loading: (Go/SD)L' = 1070(1 - 2.14 x 1O-2(LoIPL)0,,62)
Plooding:
Low Range:
(LoIPL) <. 209
(GoISD>p = 2118(1 - 2172 x 1O-2(Lolf>L)0.57) (a-Io-w)n
High Range: . (L 01 fL) > 209
(Go/SD)p = 1500(1 - 1.875 x 1O-2(LolfL)0.57) (crier w)n
n = 0.58 for (LolfL> < 209
n = 1.0 for (LolfL) > 209
Zone 1. Pre-Load and Pre-Critical:
where
(Go/SD) <. (Go/SD)L"; (LoIFL)< 209
(~P/N) = al (Gol(/)1_8
al = 1.45 x 10-6 (102.60 x 1O-3(LolfL»
Zone 2. Pre-Load and Super-Critical:
where
(Golf/) < (GOI!D)L' ; (LolfL> > 209
(~P/N) = a2 (Go/SD)le8 + Kl
a2 = 2.83 x 10-6 (102.035 x 1O-3(LolfL»
Kl = 1.38 x 10-11 (LoIPL)3. 72
E-42
-------
1-1/2-INCH CERAMIC RASCHIG RINGS Cont'd.
(3/16" WALL)
Zone 3.
Load-to- Flood and Pre - Critical~
where
(Go/'/)F >' (Go/'/) > (Go/rD)L ; (LoIfL) <: 209
(4P/N) = a3 (Go/rD)3,,0
a3 = 4040 x 10-10 (104010 x 1O-3(Lo/PL»
Zone 4.
Load-to-Flood and Super-Critical:
where
(Go/'/)F > (Go/'/) > (GO/'/)L 9 ; (Lo/PL> > 209
( t.P /N) = a4 (Go/'/)3.0 + K2
a4 = 5.77 x 10-10 (102.98 x 1O-3(Lo/PL»
K2 = 4.55 x 10-5 (Lo/fL)1.50
E-43
-------
TABLE A-4
INTRINSIC PRESSURE DROP EQUATIONS
2-INCH CERAMIC RASCHIG RINGS
Critical Gas Sweep Liquid Rate: (Lo/fL) = 240
Loading: (Go/~)L t = 1070 (1 - 2.14 x 1O-2(LolfL)0.62)
High Range:
(LolfL) < 240
(Go/9)F = 2280(1 - 2.33 x 1O-2(LoQOL)0.57)(0"'10'"w)n
(LoIPL) >" 240
(Go/~) = 2060(1 - 2.28 x 1O-~oIPL)0.57)«)'"'/ 240
Zone 1. Pre-Load and Pre-Critical:
where
(Go/~) < (Go/~)L t ; (LolfL) <. 240
(AP IN) = al (Go/~)l. 8
al = .1.00 x 10-6(103.05 x 1O-3(LoIPL»
Zone 2.
Pre-Load and Super-Critical:
where
(Go/~) < (Go/~)L' »(LoIPL) > 240
(AP/N) = a~ (Go/~)1.8 + Kl
a2 = 1.265 x 10-6(102.36 x 1O-3(Lo,fL»
Kl = 8.27 x 1O-15(LoIPL)5.3
E-44
-------
2-INCH CERAMIC RASCHIG RINGS Cont'd.
Zone 3. Load-to- Plood and Pre-Critical:
where
(Go/i/)p > (Go/f{) /' (Go/fJ)Lv ; (LoJOL> < 240
(AP/N) = a3 (Go/~)3,,0
a3 = 2055 x 10-10 (103091 x W-3(Lo/fL»
Load-to-Plood and Super-Critical:
. Zone 4.
where
(Go/'/)p > (Go/'/) > (Go/i/)L' ; ~o/PL) > 240
(AP/N) = a4 (Go/i/)3.0 + K2
a4 = 4.33 x 10-10 (l02.51 x 1O-3(Lo/PL»
K2 = 7.02 x 10-5 (Lo/PL)1.43
E-45
-------
TABLE A-5
INTRINSIC PRESSURE DROP EQUATIONS
3/4-INCH CERAMIC BERL SADDLES
Critical Gas Sweep Liquid Rate:
. Loadings:
Low Range:
High Range:
Flooding:
Low Range:
(Lo/fL) = 160
(Go/tJ)L' = 748 (1 - 2.01 x 10-2 (Lof>L)0.62).
(Go/tJ)L' = 453 (1 - 1.80 x 10-2 (LoUOL)0.62)
(L o/fL) <:. 160
(GO/tJ)F = 1825 (1 - 3.17 x 10-2 (Lo~L)0.57)(cr /crw)n
(LonoL) > 160
(GoltJ)F = 1405 (l - 2.70 x 10-2 (Lo4DL)0.57) (0" Icrw)n
0.58 for (Lo/PL) <. 160
High Range: .
n =
n = 1.0 for (LoIfL) > 160
Zone 1.
Pre - Load and Pre - Critical:
(Go/tJ) < (Go/tJ)L' j (LolfL) ~ 160
(~P IN) = al (GoltJ)1.8
al = 2.08 x 10-6 (103.25 x 1O-3(LoIPL»
where
Zone 2.
Pre-Load and Super-Critical:
(Go/f/) < (GoltJ)L t j (LotfL) > 160
(AP/N) = a2 (GoltJ)1.8 + Kl
a2 = 1.379 x 10-6 (104.68 x 1O-3(Lo/fL»
Kl = 5.77 x 10-9 (LoIPL)3.0
where
E-46
-------
3/4-INCH CERAMIC BERL SADDLES Cont'd.
Zone 3.
Load -to - Flood and Pre':' Critical:
where
(Go/'/J)p > (Go/f/» > (GO/'/)L; (LoIfL) < 160
. CAP/NY == &3 (Go/f))2075
. .
a3 . = 2071 x 10-9 (105,,607 x W~3 (Go/f) '> (Go/fD>L ; (Lo/fL) > 160
(.6P/N) = a4 (Go/'/J)2.75 + K2
a4 = 1.018 x 10-8 (102.20 x 1O-3(Lo/PL»
K2 = 1.283 x 10-5 (LoUOL)1.9
E-47
-------
TABLE A-6
INTRINSIC PRESSURE DROP EQUA'DONS
I-INCH CERAMIC BERL SADDLES
Critical Gas-Sweep Liquid Rate: (Lo/fL)c = 176.5
Loading: (Go/~)L = 1354 (1 - 2.43 x 10-2 (Lo/fL)0.552)
Flooding: (Go/0)p = 2360 (1 - 2.64 x 10-2 (Lo/PL)0.S7) (0" /(f w)n
n = 0.58 for Lo/PL ~ 176.5
n = 1.0 for Lo/fL > 176.5
Zone 1.
Pre-Load and Pre-Critical:
where
(Go/~) < (GO/0)L; (Lo/fL) L.... 176.5
(AP/N). = . a1 (Go/0)1.8
al = 1.34 x 10-6 (103.34 x 1O-3(Lo/PL»
Zone 2.
Pre-Load and Super-Critical:
where
(Go/0) < (Go/0)L; (Lo/PL) > 176.5
(AP/N) = a2 (Go/0)1.8 + Kl
, .,
a2 = 2.28 x 10-6 (102.06 x 1O-3(Lo/,oL»
Kl = 4.98 x 10-12 (Lo/PL)4.3
E-48
-------
1- INCH CERAMIC BERL SADDLES Cont'd.
Zone 3. . Load-to...P1ood and Pre-Critical:
where
(Go/'/J)p > (Go/f/» > (Go/(l))L p ~o/f>L) <: 17605
(AP/N) = ag (Go/!D)300
a3 = 2083 x 10-10 (l04Q5! :it 1O=3(Lo/~L)
Zone 4.
Load-to. Flood and Super-Critical:
where
(Go/'/»p > (Go/'/» '> (Go/'/»L; (LO/PL)
(AP /N) = a4 (Go/'/J)3.0 + K2
a4 = 5.08 x 10-10 '(103.42 x 10-3(Lo/,oL»
K2 = 1.425 x 10-5 (Lo/fL)lo67
> 176.5
E-49
-------
TABLE A-7
INTRINSIC PRESSURE DROP EQUATIONS
1-1/2-INCH CERAMIC BERL SADDLES
Critical Gas-Sweep Liquid Rate: (Lo/fL)c = 285
Loading:' (Go/~)L = 1454 (1 - 1.755 x 10-2 (LO/f>L)0.5~
Flooding: (Go/'/)F = 2740 (1 - 2.13 x 10-2 (Lo/fL)0.S1) {(f /c:rw)n
Both above equations apply to (Lo/fL) " 285
n = 0.58 for (Lo/fL) <. 285
n = 1.0 for (Lo/PL) > 285
Zone 1.
Pre - Load and Pre -Critical:
where
(Go/~) < (GO/~)L; (Lo/fOL) -<. 285
(AP/N) = al (Go/~)1.8
~ = 7.90 x 10-7 (102.13 x 1O-3(Lo/fL» .
Zone 3~ Load-to- Flood and Pre-Critical:
where
(Go/~)F > (Go/~) > {Go/~)L; (Lo/fL) .c:::. 285
(AP/N) = a3(Go/~)3.0
a3 = 1.612 x 10-10 (102.97 x 1O-3{Lo/fL»
*No data available for Zones 2 & 4 Super-Critical Operation.
E-50
-------
TABLE A-8
INTRINSIC PRESSURE DROP EQUATIONS
2-INCH CERAMIC BERL SADDLES
Critical Gas Sweep Liquid Rate: (Lo/fL> = 300
Loading: (GO/'/)L = 1692 (1 - 1017 x 1O-2(Lo/fL)0"S5)
High Range:
v...'o/ fL) -< 300
(Go/r,D)p = 3110(1 - 10866 x 1O-2(LofL)Oo57) (cr-/(J'"w)n
(Lo/fL) > 300
(Go/'/)p = 2715(1 - 1.532 x 1O-2(Lo!fL)0.57) (a-/(J'w)n
Flooding:
Low Range:
n = 0.58 for (Lo/PL) -< 300
n = 1.00 for (Lo/fL) > 300
Zone 1.
Pre-Load and Pre-Critical:
where
(Go/'/) < (Go/0)L ; (Lo/fL) L. 300
(AP/N) = al (Go/'/)1.8
a1 = 4.28 x 10-7(103.57 x 1O-3(Lo/fL»
Zone 3~
Load-to-F100d and Pre-Critical:
where
(Go/'/)F > (Go/'/) > (Go/ID)L ; (Lo/fL) <. 300
(AY/N) = a3 (Go/'/)3\>0
a3 = 7.44 x 10-11(103.46 x 1O-3(LofL»
*No data available for Zones 2 and 4 Super-Critical operation.
E-51
-------
BffiLIOGRAPHY
1)
2)
Crocker, B. B., Chem. Eng., 75, No. 15, 109-116, July 15 (1968).
Hangebrauck, R. P., et al, "Sources of Polynuclear Hydrocarbons in
the Atmosphere", fublic Health Service (1967) PB-174706.
3)
4)
, .
Hensel, S. H. & Reybal, R. E., Chem. Eng. Prog. 48, 362-370 (1952).
Howkins, J. E. & Davidson, J. F., AIChE. Journal~ 325-329 (1958).
5)
6)
Kalika, P., Chem. Eng., 76, No. 16, 133-138, July 28 (1969).
Lerner, B. J. :& Grove, C. S., Jr., Ind. Eng. Chem., 43, 216-225 (1951).
7)
Lubin, B., Ph. D. Thesis in Chemical Engineering, University of Missouri, 1949.
8)
Mickley, H. S., Chem. Eng. Prog., 45, 739-745 (1949).
9)
NASA Research, Lewis Center, "Computer Program for the Calculation of
Chemical Equilibrium Compositions" by Sanford Gordon & Bonnie McBride.
10)
Perry, J. H., "Chemical Engineers' Handbook", 4th Ed., McGraw-Hill,
New York, N. Y. (1963).
11)
Pigford, R. L. & Pyle, C., Ind. Eng. Chem. 43, 1649-1652 (1951).
12) . Prahl, W. H., Chem. Eng. 77, No. 23, 109-113, Nov. 2, (1970).
13) .
Sandomirsky, A. G., et al, "Fume Control in Rubber Processing by Direct-
Flame Incineration", J. Air. Poll. Control Assoc., 16, 673-676 (1966).
14)
Sherwood, T. K. & Pigford, R. L., "Absorption & Extraction" 2nd Ed.,
McGraw-Hill, New York, N. Y. (1952).
~5)
Treybal, R. "Mass Transfer Operations", McGraw-Hill, New York, N.Y. (1955).
16)
Wilke, C. R., Cheng, C. T., Ledesma, V. L. & Porter, J. W., Chem. Eng.
Prog. 59, No. 12, 69-75 (1963).
E-52
-------
APPENDIX F
PARTICULATE-SCRUBBING SYSTEMS
FOR
HYDROCARBON EMISSION CONTROL
-------
SUMMARY
Based on the initial finding that the major station-
ary-source emission of hydrocarbons was attributable to com-
bustion processess and the fact that these processes emitted
both carcinogenic heavy hydrocarbons and particulatess atten-
tion was focused on controls for combustion process emissions.
Review of control technology in this area showed that wet
scrubbing, although widely-useds had an inadequate design or
performance assessment base, and was characterized by an ex-
treme dichotomy between a highly-developed inertial impaction
theory and an absence of a rational engineering design basis.
In view of the literature finding that the carcinogenic poly-
nuclear hydrocarbon fractions, such as benzo(a)pyrene, emitted
by combustion operations follows the path of the particulates
in any control process, correlation of wet scrubber performance
for particulate/hydrocarbon removal was undertaken.
When dispersed-liquid wet scrubber performance data
were reviewed in terms of a mechanism of turbulent agglomeration
contacting and secondary inertial removal, rather than inertial
impaction alone, it was found that there were three primary
independent variables: inlet dust loading, water/gas ratio,
and gas velocity (or pressure drop equivalent). It was further
found that when scrubber performance was expressed in terms of
the percent mass penetrations the perf~rmance was inversely
proportional to both the inlet dust loadings Cds and the liquid/
gas ratios (WIG). When the latter variables were combined into
a single product group, termed the "agglomeration index", (Cd)
(W/G), it was found that literature data on wet scrubber dust
penetration showed penetration to be inversely proportional to
the agglomeration index. Application of this new concept to
a wide range of experimental and field data on wet scrubber
cerformance showed the agglomeration index to be a highly-
sensitive and very powerful general correlating parameter.
Deviations from the Agglomeration index correlation were found
to be indicative of dust feed pre-agglomeration or design de-
ficiencies in the removal mechanism (as distinguished from
contacting mechanism) of the scrubber.
The final primary variable of pressure drop, ~P,
was found to be readily correlated when performance data were
plotted at constant agglomeration index conditions, and the
final generalized equation obtained was:
-------
- K(~p)-n
% Penetration - TCdTTW7GT
where n = 1 or 2, depending on the flow regime, and K is the
performance constant characterizing the relative capability
of the scrubber. This ~orrelating equation allows, for the
first time, valid absolute performance comparisons between
scrubbers, and establishes the test or operating conditions
necessary to validate such comparisons.
It is apparent that established inertial impaction
theory fails to satisfactorily describe the scrubber variable
behavior found in this investigation, and there is ample reason
to doubt that this theory, in its present form, applies to
wet scrubber performance description. While it is demonstratd
that the data behavior is in accord with proposed turbuleryt
agglomeration ,mechanisms, corollary quantitative theor~tical
support is stKll to be developed. In this respect, the en-
gineering analysis of scrubber performance and the predictive
capability of the correlating equation has temporarily advanced
empiricism beyond the present re~ch of theory. '
-------
I.
II.
I I I.
I V.
V.
VI.
TABLE OF CONTENTS
INTRODUCTION
DEFINITION OF CONTROL PROBLEM
A.
B .
C.
Control Technology State-of-Art
Hydrocarbon Emission and Control Relevance
Scope of Program
WET SCRUBBER THEORY
A.
B.
C.
Inertial Impaction Theory
Turbulent Agglomeration Theory
Secondary Collection Forces
1. Diffusiophoresis and Thermophoresis
2. Condensation-Conditioning
WET SCRUBBERS: DESIGN FOR PARTICULATE REMOVAL
A.
B.
Controlling Parameters
Inlet Dust Loading
1. Data of Ingels, Shaffer and Danielson (1960)
2. Data of Johnson, et a1 (1955)
Water/Gas Ratio
Agglomeration Index
1. Data of Ingels (1960)
2. Data of Calvert and Legatski (1970)
2(a} Venturi Scrubbers: Venturi, "Ventri-
Rod" and "Ventri-Sphere"
2(b} Impingement Scrubber
2(c) "Air Tumbler"
2(d) Packed Bed Scrubber and Multiclone
2(e) "Wetted Screen"Scrubber
Particle Size Effects
1. Particle Size Distribution
2. Data of Krista1, Dennis and Si1verman(1957)
3. Size Dependent Penetration Behavior
Anomalous Behavior
1. Data of Lapp1e and Kamack (1955)
2., Data of Lancaster and Strauss (1971)
Penetration as a Function of Scrubber Pressure Loss
C .
D.
E.
F.
G.
CORRELATION IN SEARCH OF A THEORY
RESEARCH AND RECOMMENDATIONS
A.
B .
C .
Continuation of Literature Data Correlation
Development of a Satisfactory Theoretical Model
Experimental Program
NOMENCLATURE
BIBLIOGRAPHY
Page No.
F-l
F- 3
F- 3
F- 4
F-5
F-6
F-7
F-ll
F-15
F-15
F-16
F-18
F-18
F-19
F-20
F-27
F- 31
F- 32
F-34
F-38
F-38
F- 44
F-48
F-52
F-62
F-65
F-65
F-67
F-73
F-78
F-78
F-81
F-81
F-91
F-94
F-94
F-94
F-95
F-97
F-99
-------
Table No.
4-1
4-2
4-3
4-4
4-5
4-6
4-7
4-8
4-9
4-10
4-11
4-12A
4-12B
4-13
LIST OF TABLES
Page No.
Test Data from Asphaltic Concrete Plants
Controlled by Scrubbers
Fractional Collection Efficiency Data for
Scrubbers Serving Asphaltic Concrete Plants
F-23
F-26
Gross Penetration for a Venturi Scrubber for
an Inlet Dust of Mean Diameter 1.35 Microns
and Standard Deviation 1.7
F - 39
Gross Penetration for a "Ventri-Rod" Scrubber
for a Coal Dust of Mean Diameter 1.42 Microns
and Standard Deviation 1.64
F-41
Gross Penetration for a "Ventri-Sphere"
Scrubber for an Inlet Dust of Mean Diameter
1.44 Microns and Standard Deviation 1.7
F-45
Gross Penetration for an Impingement Scrubber
for an Inlet Dust of Mean Diameter 1.44 Microns
and Standard Deviation 1.7
F-49
Gross Penetration for an "Air Tumb1er" for an
Inlet Dust of Mean Diameter 1.35 Microns and
Standard Deviation 1.7
F-53
Gross Penetration for a Dry Packed Bed Scrubber
for an Inlet Dust of Mean Diameter 1.82 Microns
and Standard Deviation 1.68
F-57
Gross Penetration for a Wet Packed Bed Scrubber
for an Inlet Dust of Mean Diameter 1.44 Microns
and Standard Deviation 1.74
F-58
Gross Penetration for a Mu1tic1one for an Inlet
Dust of Mean Diameter 1.5 Microns and Standard
Deviation 1.7
F-59
Gross Penetration for the "Wetted Screen"
Scrubber for an Inlet Dust of Mean Diameter 1.5
Microns and Standard Deviation 1.8
F-63
Multi-Venturi (Solivore) Scrubber Penetration
Test Data for Three Aerosols at Varying Water Rates
Penetration Test Data for Varying Inlet Dust
Loading
F-68
F-68
Effect of Sampling Method on Estimation of
Effluent Loading Single Stage Solivore Scrubber
F-72
-------
Figure No.
4-1
4-2
4-3
4-4
4-5
4-6
4-7
4-8
4-9
4-10
4-11
4-12
4-13
LIST OF FIGURES
Wet Scrubber Efficiency Asphaltic Concrete
Plant Dust. .
Penetration as Function of Scrubber Inlet
Dust Loading Hot Mix Asphalt Paving Plants
Penetration Dry vs. Wet Cyclone Operation
as Function of Inlet Dust Loading
Penetration Difference for Dry vs. Wet
Cyclone Operation as Function or-Inlet
Dust Loading
Penetration as a Function of Agglomeration
Index, (Cd)(W/G) - Multiple Centrifugal
Sc~ubbers - Oil Fired Dryer
Penetration as Function of Agglomeration
Parameter, (Cd)(W/G) - Multiple Centrifugal
Scrubbers - Gas Fired Dryer
Penetration as Function of Agglomeration
Index, (Cd)(W/G) - Baffled Tower Scrubbers
Venturi and "Ventri-Rod" Scrubbers Gross
Penetration as Function of Agglomeration
Index, (Cd) (W/G)
"Ventri-Sphere" Scrubber Gross Penetration
as Function of Agglomeration Index,(Cd)(W/G)
2-Stage Impingement Scrubber Gross Penetration
as Function of Agglomeration Index,(Cd)(W/G)
"Air Tumbler" Gross Penetration as Function
of Inlet Dust Loading
Dry and Wet Packed Bed Scrubber and Multiclone
Gross Penetration as Function of Agglomeration
Index or Dust Loading
Wetted Screen Scrubber Gross Penetration as
Function of Agglomeration Index, (Cd)(W/G)
Page No.
F-22
F-24
F-28
F-30
F-35
F- 36
F-37
F-42
F-47
F-51
F-56
F-60
F-64
-------
Figure No.
4-14
4-15
4-16
4-17
4-18
4-19
4-20
5-1
LIST OF FIGURES (Continued)
Solivore Scrubber Penetration as Function
of Agglomeration Index, (Cd)(W/G)
Solivore Scrubber Penetration as Function
6f Agglomeration Index, (Cd)(W/G):Total Data
Steady~State Aerosol Mass Concentration vs.
Minimum Particle Size --
Performance of l-Inch S-Bend Contactor and
8-Inch .Cyclonein .Seri~s Penetration as
Function of Agglomeration Index
Typical Behavior of Penetration as a Function
of Scrubber Pressure Drop
Penetration as a Function of Pressure Drop:
Odor Removal
Ventuir Removal of Phosphoric Acid Penetration
as Function of Pressure Drop
Log-log Plot o~ Penetration Data of Brink
and Contant
Page No.
F-69
F..70
F-75
F-80
F-82
F-85
F-88
F-92
-------
L
INTRODUCTION
The purpose of this study was to identify and eval-
uate control techniques for the prevention or removal of
hydrocarbon emissions from stationary sources, with the ul-
timate objective of identifying gaps in methodology or design
procedures and formulating remedial experimental programs.
A review of the open literature on hydrocarbon emission con-
trol methods showed that, aside from the compilation, Air Po1-
~~ti~~ ;~;i~~e~~g~i~~~~a~u~~~~iai~~~)in~~~~~i~~ ~~;{i~~i~~~
information for control processes or equipment. A general
review of the background control art has recently been pre-
sented by HEW (1970) in their manual, Control Techniques for
Hydrocarbons and Organic Solvent Emissions from Stationary
Sources, but this treatment is largely qualitative.
The crux of the control evaluation problem was iden-
tified in a recently-completed specific study of heavy hydro-
carbon emission. In their preliminary report on Control Tech-
ni ues for Pol c clic Or anic Matter Emissions, GCA Technology
1970 , after considering the available quanitative design
data, stated, "in the absence therefore of any practical in-
formation on the effects of existing types of gas cleaning
equipment on polycyclic organic emissions, we must rely on
theoretical considerations to determine which cleaning tech-
niques will be effective in removing polycyclic organic emis-
s ion s II . W h i 1 e the s tat e - a f - the-art out 1 a a k for. can t r a 1 tech -
no10gy for certain hydrocarbon classes other than polycyclic
hydrocarbons :is not quite this bleak, the degree of control
effectiveness offered by a specific design for a specific
emission is still uncertain in the best of cases.
Control effectiveness uncertainties may result in
many instances from a variable definition of pollution, or
fai1ure'to measure the process output under operative controls.
An example is provided by the review of Fawcett (1970) .of the
emissions common to phthalic anhydride manufacturing industry,
and the nature of the controls used. The emissions are the
acids or acid anhydrides, aldehydes, a-xylene or naphthalene,
and occasionally, particulates. In one set of locations, the
air pollution problem is that of notsance odors and lachrymatory
compounds, which are characteristic of the organic acid com-
ponents. On the other hand, in West Coast locations, or others
F-1
-------
with poor air shed ventilation, the pollution problem is that
caused by the photochemically-reactive smog-forming aldehyde
and aromatic constituents. Wet scrubbing is reported by
Fawcett to afford 99% removal of the acids and anhydrides, but
to give relatively poor control of the aldehydes. Thus, in
smog-prone areas where the aldehyde is the primary concern,
"control" would involve incineration of the process off-gas,
either by direct-flame or catalytic means, while "control" in
other areas would call for acid removal by wet scrubbing.
Control effectiveness is therefore measurable only in the
context of a variable local pollution problem. With a limited
number of exceptions, the health and pollution hazards of
specific organic compounds have not been fully determined,
and until these basic parameters are defined, control effective-
ness cannot be absolutely fixed. Thus, quantitative design
and II e f f i c i e n c y II i n form a t ion for e xis tin g coli t r 0 1 tee h n i que s
can only be assembled on a "case-history" basis. This type
of information can be obtained by (a) soliciting information
on successful field installations, (b) utilization of unpub-
lished information from the files of engineering firms in
the emission control field, or (c)development of original de-
sig~ techniques by organization and evaluation of published
experimental and field data in a defined control area. The
latter two avenues were the ones chosen for implementation
in this program.
F-2
-------
II.
DEFINITION OF CONTROL PROBLEM
A.
CONTROL TECHNOLOGY STATE-OF-THE-ART
The starting point of this program was the public-
ation, Control Techni ues for H drocarbon and Or anic Solvent
Emissions from Stationary Sources HEW, 1970. A review of
this compilation and supporting literature indicated that ap-
plicable technology for the control of hydrocarbon emissions
was very unevenly developed. Design methods for techniques
such as adsorption and after-burning were available and reason-
ably straightforward, whereas wet scrubbing methods, includ-
ing absorption, direct-contact condensation, and wet partic-
ulate removal, appeared to lack an adequate design base. This
conclusion was partially confirmed by the statement regarding
particulate collectors made in the recent control equipment
review by the American Industrial Hygiene Association (AIHA,
1968) which states, "when the process is on the drawing board
and the collector is to be specified as part of the plant de-
sign, the prime recourse is to experience on similar plants
and processes". The frequent association of higher molecular-
weight hazardous organic fractions with particulate emissions
from combusti6n" processes makes this apparent design deficiency
of more than academic interest.
Despite the fact that wet scrubbing methods are
widely used for particulate and hydrocarbon removal, usually
concurrently, an assessment of the present availability of
design methods and the general state of technical development
of this area showed there to be a very serious deficiency in
technology. While the nature of this deficiency is examined
in detail in Section 3 of this report, a technical literature
review showed that there was an unbridged gap between theory
and field practice, and that there was no rationale for scrubber
selection or design for the following problem areas:
( a )
Removal of liquid hydrocarbon aerosols or of
particulates in association with heavy hydro-
carbon co-particles, particular]y polynuclear
h.ydrocarbons.
(b)
Combustion gas scrubbing with secondary ab-
sorption of water-soluble oxidized hydro~ar-
bons.
F-3
-------
Condensation (direct',or indirect) followed
by mist-filtration or scrubbing.
Thus, there was an obvious necessity for an intensive
critical examination of scrubber technology and a clear mandate
for re-evaluation of design approaches and data corre1ati~n.
This conclusion was supported by a preliminary review of the
relative seriousness of hydrocarbon,emissions from stationary
sources.
( c)
B.
HYDROCARBON EMISSION AND CONTROL RELEVANCE
At the very outset of thns systems study, it,beca~e
apparent that the major source of stationary hydrocarbon emis-
sion was the combustion process. Invariably, such processes
emit particulates in association with hydrocarbons. Power
generation, the incineration of municipal wastes, and the dis-
posal of industrial organic liquid sludges are examples of such
operations. A good deal of attention has been foc~sed on, this
particular aspect of hydrocarbon-particulate emission because
of the indicated health problem attendant on the polycyclic
aromatic hydrocarbons normally found in combustion off-gases,
particularly in inefficient combustor units. Several of the
polycyclic nuclear a~omatics~ notably benzo(a)pyrene (BaP),
are known to be carcinogenic, and for this reason, control
information for these hydrocarbons merits priority.
Hangebrauck,'(1967) measured rates of emission of ,
polynuclear hydrocarbons for refuse burning and various in-
dustrial oper~tions, both before and after the exhaust gases
passed throu~h control equipment. Data were presented on the
effect of CO waste heat boilers and plume burners on catalytic
cracking regenerators in petroleum refining, and of water- '
spray scrubbers and steam "spr~ys" in batch asphalt hot-road-
mix plants. In the latter case, water-spraying was found to
be 92%, effective in BaP removal, but the data were extremely
l1mited, and. no information was given on the design of,the
contro~ equipment.
Cuffe (1967) studied the emission of seven key
polynuclear hydrocarbons from coal-fired power plants using
various boiler designs. The most significant finding was'
the fact that preliminary tests showed that considerable
recovery of the polynuclear aromatics was effected by the
fly-ash collectors, although no quaTtitative data' were pro-
vided. Full-scale tests showed that even forma1dehyde'fol-'
lowed the particulates, indicating probable adsorption. In
terms of control relevance, this means that hydrocarbon emis-
.F-4
-------
sion control for a combustion process can and should be ac-
complished via particulate control. This restatement of
control objective, along with the previously established
design gap for wet scrubbers, served to define the priority
of control technology assessment as being that of partic-
ulate removal by wet scrubbing techni~ues.
c.
SCOPE OF PROGRAM
Initially, the study of wet scrubbing methods for
the control of hydrocarbon emissions from stationary sources
was considered to be the first phase of an overall control.
method design workup. However, preliminary efforts on wet
scrubber particulate removal data correlation proved 50 pro-
ductive that work was continued on this subject to the limit
of the project time. As a complementary study, material was
assembled on the design of packed scrubbers for incinerator/
scrubber trains in cooperation with the engineering staff of
the Maurice A. Knight Company of Akron, Ohio. These new and
hitherto proprietary design techniques have been presented in
a corollary report covering methods for estimating the removal
of soluble pollutants such as oxidized hydrocarbons, in a
combustion gas quench-cooling operation. A third effort,
covering proprietary condensation-mist elimination techniques
for controlling certain hydrocarbon solvent emissions, was
abandoned when permission to re1ease the data was withdrawn
by the cooperating engineering firm.
F-5
-------
I 11.
WET SCRUBBER THEORY
Despite the fact that water-scrubbing of combustion
gases or other emissions is one of the oldest techniques em-
ployed for particulate removal, the contacting mechanisms in-
volved are still not well understood. Consequ~~tly, as Semrau
(1963) pointed out, "failures have resulted from ignorance of
the factors involved in scrubber performance and from the, lack
of a rational basis for design". A cursory review of the cur-
rent available literature indicates that this situation still
prevails.
A number of recently-published reviews and books
have contained sections or chapters ostensibly dealing with
wet scrubber design, wherein very little, if any, quantitative
design information is presentedJ HEW AP-51 (1969), Control
TechniQue~ for Particulate Air Pollutants, presents. an excel-
lent concise review of the various types of wet scrubbing
equipment, along with information on the operating and per-,
formance ranges of each type. However, although the mechanisms
of contacting and operation of the different scrubber types
are briefly explored, no generalized design criteria are cited.
Similarly, in Chapter 4 of Danielson (1967), dealing with con-
trol equipment for particulate matter, the section on wet
scrubbers (by E.J. Vincent) contains only qualitative descriptive
matter, with no menti~n of controlling parameters. The same
chapter, however, deals fully, and. in detail, with des.ign ap-
proaches for other equipment such as inertial separators (dry),
fibrous filters, and electrostatic units. The AIHA volume of
the Air Pollution Manual covering control equipment (Part II,
1968) also confines its discussion of wet scrubbers to descriptive
material, augmenting this with Stairmand,'s (1956) "typical"
grade-efficiency curves for wet scrubbers.
At least part of the present unsatisfactory state
of development of wet scrubber design art may be attributed
to failure to recognize the dual-functionality of the wet
scrubber. Thus, Lunde and Lapple (1958) in their state-of-
the-art review, speak only of "deposition" mechanisms in wet
scrubbers, indicating these to be solely surface collection
and removal processes. Only recently, in work such as that
of Davis and Truitt (1971), has attention been devoted to the
other fundamental scrubber mechanism: that of particle growth
by agglomeration. Most wet scrubbers utilize some type of
F-6
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inertial separation force to cause deposition of particles of
a minimum size or mass on a surface collector, but the scrubbers
also effect particle "growth" to the minimum removal size by
contact between water droplets and dust particles. There is
also evidence that particle growth by collision processes oc-
curs even for dry centrifugal collectors such as cyclones, and
indeed, in the field of atmospheric dust fallout studies, con-
siderable attention has long been given to agglomerative
particle growth. The theoretical and experimental work of
Langstroth and Gillespie (1947) is of fundamental significance
in this latter area, and would appear to have immediate rele-
vance to the interpretation of wet scrubber performance data,
and will be considered later in this section.
The inability to isolate the primary contacting mech-
anisms has resulted in a highly uneven theoretical approach to
wet scrubber theory, and the greatest amount of development
effort has been devoted to inertial impaction theory and the
least attention to turbulent agglomeration theory. Neither
theory has been directly and fully tested against performance,
and recent sophisticated attempts to compare inertial impaction
theory with scrubber performance, such as the computer-as-
sisted experimental investigation conducted by Boll (1971)
have proved disappointing. Because the impaction theory ap-
pears to have merited the major share of attention and de-
velopment, it will be considered first.
A.
INERTIAL IMPACTION THEORY
The inertial impaction theory, in its quantitative
form, was initially set forth Langmuir and Blodgett (1944),
and is essentially a "ballisticll collision theory. It is
generally assumed that the small dust particulates move with
the velocity of the gas stream, and that the larger, slower-
moving water droplets serve as collision IItargetsll for the
dust particles. These concepts and elementary equations for
the volume and number of particulates swept by a water drop
swarm in a scrubber had been considered earlier by Kleinschmidt
(1939), but it was assumed that all dust particles in the
volume swept by the drops would be captured. Langmuir.and
Blodgett pointed out that small particles of low enough in-
ertia 'could follow the flow streamlines around the drop, and
introduced the concept of "target efficiency", which is the
ratio of the number of particles actually captured by the'
drop to:the total number in the gas volume swept by the drop.
F -7.
-------
Target efficiency has been defined in terms of
various combinations of its controlling variables, but the
parameter generally used in the United States is the one
suggested by Langmuir and Blodgett (1944):
'I'=~
~
( 3-1 )
where
'I'
= separation number or inertial impaction
parameter', dimensionless
Dd = diameter of water droplet
Dp = diameter of dust particle
V9,Vd = ve~ocities of gas and water drop respectively
Pp = density of dust particle
J.I
= viscosity of gas
Because target efficiency for drop collectors can be theo-
retically calculated as a function of the separation number,
'1', only for viscous flow (NRe103)
ranges with respect to the drops, the relation between tar-
get efficiency, Et, and'!' is best determined experimentally.
Experimental results for rigid spheres have been presented
by Ranz and Wong (1952) and for supported liquid drops by
Walton and Woolcock (1960). However, as the review of these
and other data by Boll (1971) indicates, agreement is only
fair. The use of target efficiency vs. inertial parameter
plots for calculation of spray chamber contact times for a
given dust and water drop size has been demonstrated by Ranz
and Wong (1952) who stated the time rate of aerosol concen-
tration change in a uniform spray region as:
-dNp/dt = NpNdEtVgnDa/4
(3-2)
where
Np = number concentration of particles, cm-3
Nd = number concentration of drops, cm-3
Vg = relative velocity of droplets through the
dust, cm/sec.
Dd = diameter of water droplet, cm.
F-8
-------
Integration of Equation (3-2) over a finite holdup time, t,
yields a fractional residual, or weight penetration as:
F = Np/Npo = exp(-tVgNdEt~0~/4) (3=3)
Equation (3~3) states that scrubber dust penetration should
yield a straight line on a semi-log plot against any of the
variables in the right-hand parenthetical group, or any pri-
mary operating variable which is a direct function of these
quantities. Variants of Equation (3-3) have been used in ap-
plication of inertial impaction theory to scrubber efficiency
data, and Johnstone, Feild and Tassler (1954), assuming Et
was proportional to ('1') 1/2, correlatEi'd Venturi data by the
relation:
1/2
E = 1 - exp[-K(W/G)('1') ]
( 3- 4)
where
(W/G) = liquid/gas ratio, gal./MCF
However, examination of the "straight-line" plots of F vs.
(W/G)('1')1/2 shoWs that the experimental data collected by
these authors do not quite follow..the predicted line; the
separate particle run series all show regular concave-up-
ward curvatOre. Earlier, Johnstone and Roberts (1949) had
correlated Venturi penetration data for dust collection,
502 absorption, and humidification on the basis of the
specific drop area exposed per cubic foot of gas, without
recourse to inertial impaction theory. Similarly incon-
sistent experimental da~a with respect to impaction theory
are still being obtained; Boll (1971) found a fair fit of
impactiQn theory to Venturi scrubbing data at low pressure
drops of about 30" W.C., but virtually no correlation for
higher ~p operation.
One of the basic difficulties of Equation (3-1)
is that it properly applies to uniform particle and drop
sizes, and appncation to practical systems with a spectrum
of sizes presents some difficulties. If the water drop
size is taken to be uniform, then the overall efficiency
is the summation of the target efficiency of each partitle
size multiplied by the fraction of the particles in that
particular size range. If efficiency is defined as the
weight fraction removed, E, then:
E = (02 E t ( d M / d 0 p) d 0 P
)01
( 3- 5)
F-9
-------
where (dM/dDo) is the slope of the cumulative size distripution
curve based on mass fraction analysis (Calvert and Taheri,
1966). While Equation (3-5) may account for particle size
variation, it is valid only for a single water drop size, and
the normal size distribution of water drops produced by spray
nozzles, or by atomization within equipment presents a problem
of extrapolation. In addition, there are fundamental difficulties
deriving from target efficiency dependence on physical system
properties which do not appear in Equation (3-1). Goldshmid
and Calvert (1963) found that target efficiency was a function
of the degree of mutual wetting (interfacial tension) of the
aerosol and liquid pair. Further, wake capture behind the
liquid drops was found to significantly increase the experi-
mental target efficiencies at high air velocities and small
particle sizes. High air velocities give rise to drop in-
stabi"lity, and above approximately 200 ft/sec relative gas
velocity, drops will oscillate and shatter. Drop atomization
due to gas shear is usually handled by the Nukiyama and
Tanasawa (1938) empirical correlation which, for the general
case, may be stated as:
585«1L) 1/2
Do = V ( ) 1 f 2
9 PL "
+ 59 7 [ II / ((] P ) 1/2 ] 0 . 45 ( 1 000 L iG ) 1. 5
"L L L
(3-6)
where
Do
= Sauter surface-to-volume mean diameter, microns
L/G = ratio of volume flow of liquid to volume flow
of gas at vena contracta
Vg
(]L
= relative gas velocity, meters/see
= liquid surface tension, dynes/em
= liquid viscosity, poises
llL
PL
= liquid density, g/cc
The Nukiyama-Tanasawa equation has been one of the
primary tools used in exploring the function of the high-en-
ergy loss Venturi scrubber which, like most wet scrubber units,
was an empirical development for particulate-removal service.
A good deal of attention has been devoted to exploration of
the mechanics of particulate contact in the Venturi wet scrubber,
and the derivation of Winklepleck (1970), introducing the pres-
sure~loss term into equation (3~3) is of interest. The transfer
of energy from gas to the liquid by frictional drag forces can
F-10
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be measured as the rate of gas pressure change with respect
to time, and can be expressed in terms of a drag coefficient,
Cf' so that
-dP/dt = (Stokes' drag force) x (droplet velocity
relative to gas) x (droplet concentration)
-dP/dt = [(Cf/2g)(~D~/4)(pgVg)] (Vg)(Nd)
Combining Equation (3-7) with (3-2) yields
-2Etg
-dNp/Np = CfPgV~ dP
(3-7)
(3-8)
Integration of Equation (3-8) with the boundary condition
that at zero pressure drop the outlet particle concentration
equals t~e inlet concentration' yields:
~2Etg~P
1n[(Np)out/(Np)in] = CfPgV~
(3-9) ,
Recogni~ing that the left-hand side of Equation (3-9) is
simply 1n(F), then
F = exp(-2Etg ~P/CfPgV~) (3-10)
Thus, Equation (3~10) calls for a semi-log relation-
ship between weight percent penetration and pressure drop, and
the form of the equation is similar to the relationship of
Equation (3-4) for the variable, (W/G). Just as appears to
be the c~se for the actual experimental data on the effect of
(W/G), the literature test data for the relationship between
F and ~p do not follow the indicated semi-log relationship.
Detailed comparisons will be made in the next section of this
report, but an examin~ti6n" of alternative theoretical bases
is in order.
B.
TURBULENT AGGLOMERATION THEORY
Contact between gas-suspended particles and/or drops
can occur by various mechanisms other than simple ballistic
impact induced by a large differential velocity. Ordinary
Brownian motion can induce coagulation of parti~les, but it
can be readily shown that the time and concentration require-
ments for a significant contribution from this source is out-
side the range covered by most wet scrubber applications.
For the mechanisms of localized velocity gradIent contacting
F-ll
-------
prevalent in turbulent flow, the published viewpoints are con-
flicting. For example, Lancaster and Strauss (197la) examine
the case for the turbulent velocity gradient for a duct velocity
of less than 40 ft/sec and conclude that turbulent agglomer-
ative contact would not be significant as compared to Brownian
coagulation. However, a few pages later, these same authors
note the work of Levich (1962) to the effect that lithe dis-
tortion of stream lines due to intense turbulence dominates
the distortion due to inertial effects and ... each encounter
based on straight line trajectories would be realized". There
has been an almost intuitive recognition of this latter point
on the part of a number of investigators of scrubber performance.
For example, Kristal, Dennis and Silverman (1957) applied in-
ertial impaction theory to their wet scrubber data, but quali-
fied their approach by taking droplet projection velocities
based on turbulent flow (despite a drop NR in the 35 to 335
range) citing Dalla Valle's statement (19~8) "If the fluid is
in turbulent motion, the motion of a particle injected into it
will be.turbulent regardless of the relative velocity between
the particle and the fluid". .
Despite the definite thread of intuitive acceptance
of turbulent mixing as a primary contacting mechanism for wet
scrubbers, no corresponding theoretical development has been
applied to such equipment. The only turbulent contact theory
for particle agglomeration in the literature appears to be
that of Langstroth and Gillespie (1947) and this was derived
for the case of self-agglomeration of smokes. Nevertheless,
the development is believed of direct interest to wet scrubber
data interpretation.
For the case of an aerosol suspended in turbulent
air, Langstroth and Gillespie distinguish between the mass
concentration, M , and the particle concentration, N , a
distinction not Directly made in inertial impaction ~heory.
The rate of loss of mass to surfaces contacted is taken as
proportional to the mass concentration, so that:
dMp/dt = -a Mp
where a is the "mass loss constant".
of time, and if Mp = (Mp)o at t = 0,
( 3- 11 )
If a is independent
then
In[Mp/(M~)O] = -a t (3-12)
The experimental data obtained validated Equation (3-12) and
it should be noted that this equation states a semi-log re-
F-12
-------
1ation between penetration and residence time. For a fixed
time interval, mass penetration would be inversely proportional
to the surface area~ai1ab1e for contacting, and for the case
of the wet scrubber where dispersed drops provide the surface
loss area, the drop surface area per unit gas volume would be
the primary variable, as was found by Johnstone and Roberts
(1949) for their Venturi data.
If the number~ rather than the mass of particles,
is now considered, the change of a particle striking a sur-
face is proportional to the number concentration, Np' so that
surface loss is
(dNp/dt)s = B Np
( 3- 13)
However, particle concentration can decrease by partic1e-
particle collision as well as particle-surface collision, and
since the rate of collision between two particles is propor-
tional to Np' the collision loss is
(dNp/dt)c = -kNp (3-14)
and the total rate of decrease of particle concentration is
given by:
Integration
dNp/dt = -(kNp2 + BNp)
oJ Equation (3-15) yields
1n(1/Np + k/B) = 1n[1/(Np)0 + k/B)
(3-15)
+ Bt
(3-16)
By making independent measurements of mass and particle number
loss with time, Langstroth and Gillespie were able t4 evalu-
ate the separate mass and number loss constants, a, B, and k,
and determine their individual behavior with increasing air
motion.
For a homogeneous, aerosol, the ratio, (aIB) should
be 1.0, while for aerosols having the usual particle size
distribution, the ratio was found to be grea.ter than 1.0.
This was explained on the basis that the loss of the heavier
particles was favored by the gravitational force which was
operative over the appreciable time intervals employed, and
by inertial forces resulting from turbulent deflection of the
air currents at the collecting surface of the containing box.
The relative importance of the agglomeration and surface loss
mechanisms would be expected to depend strongly on the conditions
F-13
-------
used and the particle concentration. Inspection of the re-
spective loss equations shows that for high particle number
concentrations or high dust loadings, agglomeration which
depends on (N )2, would be favored, while the rate of sur-
face loss, degendent on N , would dominate for relatively
low dust loadings and/or Reavier particles. For a scrubber
or other particulate removal device, the surface loss or
removal rate would be governed by the magnitude of the re-
moval force designed into the unit: this is almost always
an inertial removal force induced by deflection or reversal
of gas flow, or by centrifugal action. On the other hand,
for a wet scrubber the mechanism of particle growth to re-
moval size would be both that of turbulent contact between
dust and water drops, and dust-dust agglomeration. It is
doubtful if the usual fixed-drop target efficiency impaction
data is relevant to the random turbulent-contact case.
It is interesting to note that the recent theoretical
work of Davis and Truitt (1971) along similar lines to those
taken by Langstroth and Gillespie, very closely predicts the
particle size distribution vs. time curves experimentally de-
termined by these latter investigators, although no reference
is made to this fact. The infOrmation contained in these
studies of atmospheric aerosol removal phenomena has not been
utilized in analyzing wet scrubber behavior, but would seem
to offer a highly promising approach.
There is, at present, the inevitable number of con-
tradictory statements concerning turbulent agglomeration
which must be reconciled by further work. For example, Davis
and Truitt state that "agglomeration by turbulence is pre-
dicted to occur at rates proportional to the cube of the
particle diameter, hence larger particles are more subject
to growth in a turbulent mediumll. This concept is in direct
opposition to Equation (3-14) which states that collision
growth is more likely for the smaller particles simply be-
cause there are a greater number of them.
In direct scrubber work, the observation that gas
turbulence appeared to control the removal efficiency of wet
scrubbers was made by Lapple and Kamack (1955) on the basis
of extensive test work. This observation was later expanded
by Semrau (1960) into the "contacting-powerll concept, which
states that the number of contacting transfer units is pro-
portional to an exponential function of the tota1 power input
to the unit. Aside from the philosophical criticism that
transfer unit theory should' be properly applied only to a
differential contacting unit, which wet scrubbers do not ap-
proximate, the use of contactin"g power represents a drastic
F-14
-------
oversimplificaticon -of the original Lapple and Kamack data.
Further, the use Qf the contacting-power approach has not
resulted in any ability to compare scrubber performance on
either an absolute or relative basis, and the slopes of the
log-log NTU vs. power input plots appear to vary randomly
(Semrau, 1 96OT. In fact, close examination of the original
data used for such plots shows that the method of p~btting
often conceals significant and real relationships between
omitted parameters and scrubber efficiency. However, the
basic premise that turbulence is the primary mechanism of
contact-removal is consistent with most data on scrubber
performance, and the deficiency of the contacting power
correlation lies in its form, not in its premise. A directed
effort at further development of turbulent agglomeration
mechanism elucidation and theory is obviously required;
this approach appears to have been neglected because of the
possibly premature acceptance of impaction theory as the
valid theoretical framework.
C.
SECONDARY COLLECTION FORCES
1.
Diffusiophoresis and Thermophoresis
There are a number of fundamental physical pro-
cesses which assist in driving a small particle toward a
collecting surface, either drop or wall, but which are not
generally considered as primary collection mechanisms.
These include:
- ( a)
(b)
Diffusiophoresis, also called sweep
diffusion, or Stefan flow, which re-
sults from the flux of diffusing
molecules under a concentration grad-
ient. For example, when water vapor
is condensing on a surface, the water
molecules diffusing or sweeping toward
that surface will tend to move particles
along with them.
Thermophoresis which results from the
more rapid molecular motion at higher
temperatures than at low, so that a
particle in a temperture gradient is
struck more frequently on the "hot"
side and is driven toward the lower
temperature level.
F-15
-------
Diffusiophoresis and thermophoresis theory have been recently
and thoroughly reviewed by Davis and Truitt (1971) and Lan-
caster and Strauss (1971a), and this material need not be re-
peated here. Attempts to employ diffusiophoresis as an
auxiliary collecting mechanism usually involves steam injection
coupled with subsequent condensation on a cold surface or
water spray. Such a procedure not only involves diffusio-
phoresis, and thermophoresis as well, but direct condensation
on at least a fraction of the particles, causing particle
growth. The sharply-improved particulate collection efficiencies
for certain humidification-quenching sequences must be attri-
buted to condensation particle growth because the magnitude
of the change is too large to be accounted for by the usual
secondary collection mechanisms. Condensation is not truly a
collection mechanism, b'ut rather a particle-conditioning pro-
cess, but the diffusiophoretic and thermophoretic effects are
an inextricable part of any such vapor condensation operation.
2.
Condensation-Conditioning
Fahnoe (1951), in one of the earliest studies of con-
densation effects, investigated the buildup of sodium chloride
aerosol particles by steam condensation, and found that con-
siderable size growth could be achieved. For both a cyclone
scrubber and a Peabody impingement plate unit, it was deter-
mined that the overall collection efficiency of the aerosol
increased from 40 to 90% for an injection rate of 0.01 1b steam/
CF air processed. Surprisingly, it was found that only 10% of
the steam condensed on the particles, with the remainder either
condensing on the vessel walls or undergoing homogeneous nu-
cleation. This distribution of the input steam raises a number
of questions which, despite the recent research activity in
t his area, h a ve not bee n sat i s fa c tor i 1 y an s w ere d . For e x amp 1 e ,
the collection efficiency improvement has been attributed to
the nucleation of the fog directly on the particulates, which
is the usual meaning of particle growth. However, homogeneous
n u c 1 eat ion. w 0 u 1 d con t rib ute an e f f e c t =i ve i n c rea s e i n par tic 1 e
population, and examination of the Langstroth-Gi11espie state-
ment of Equation (3-14) certainly indicates that the incre-
mental particle concentration would be an additive effect.
While all other effects of steam condensation have been con-
. sidered and investigated in recent investigations, the "POp-
ulation exp10sion" caused by homogeneous nucleation of fog
particles has not been examined.
Recent work by Lancaster and Strauss (1971b) on
the effect of steam injection in wet scrubbers showed that
the improvement in scrubber performance was determined by
the quantity of steam injected, and not by the quantity con-
F-16
-------
densed on the particles. This is in direct contrast to the
careful experimental work of Davis and Truitt (1971) on a
simple condensation tube, who found that particle removal was
proportional to the steam condensation rate. Davis and Truitt
further found a marked difference in particle growth behavior
between hydrophilic and hydrophobic aerosol particles. In
a condensation field, hydrophilic and hygroscopic particles
all showed fairly uniform growth; conversely, the hydrophobic
particles showed massive growth for only a small fraction of
the total particles, with the remainder showing no growth at
all. The simpliCity of the test equipment allowed a dis-
tinction to be made between growth and removal processes, and
perhaps the most important finding of this work was that the
gas Reynolds number was a determinant of the efficiency of
removal, and this effect was attributed directly to turbulent
agglomeration.
F-17
-------
I V .
WET SCRUBBERS:
DESIGN FOR PARTICULATE REMOVAL
A.
CONTROLLING PARAMETERS
The variables governing wet scrubber performance
for particulate removal have been investigated by Lapp1e and
Kamack (1955), Johnson et a1 (1955), Krista1, Dennis and
Silverman (1957), Brink and Contant (1958), Silverman and
Davidson (1956), and Ingels, Shaffer and Danielson (1960).
Most recently, Calvert and Legatski (1970) reported on an
extensive investigation of seven commercial types of wet
~crubbers, and Lancaster and Strauss (1971) presented data
on the effects of steam injection on collection efficiency.
Although these investigations covered a wide variety of wet
scrubbing equipment, both low and high-energy types, there
appears to be general, if not universal, agreement that
there are three primary variables governing performance:
( a )
(b)
(c)
Inlet dust loading
Water/Gas Ratio
Gas pressure drop and/or gas velocity
Most of the cited literature presents performance
correlations in terms of efficiency of removal as functions
of each separate primary variable, although in many instances,
maintenance of constancy of the other two primary variables
is not followed. In addition to the three primary variables
indicated above, data are occasionally provided on the effects
of a group of secondary variables, including:
( a)
(b)
( c)
(d)
Inlet dust particle size and distribution
Scrubber type
Water droplet size and injection velocity
Nozzle location
In all cases of dispersed-liquid capture of partic-
ulates, examination of the experimental performance data shows
that, for a given dust and given set of secondary variables,
F-18
-------
it is possible to completely describe the efficiency of a
scrubber solely on the basis of the three primary variables.
Failure to recognize this fundamental relationship may give
rise to a considerable degtee of misinterpretation of data
and invalid conclusions concerning relative unit performance.
In fact, the present absence of a useful basis for comparison
of equipment performance may be attributed to this failure,
and one of the primary objectives of this data review is to
remedy this deficiency.
B .
INLET DUST LOADING
The inlet dust loading variable has the distinction
of being ignored in most theoretical inertial impaction treat-
ments of the scrubbing mechanism, but requiring almost immedi-
ate experimental attention in actual test work because of its
dominant effect. Even after its laboratory or test identification
as a primary, if not controlling variable, it may later be
overlooked in secondary efforts at correlation. For example,
in the work of Lapp1e and Kamack (1955) on several types of
scrubbers, the inlet dust loading was the initial parameter
investigated, and because of its profound influence on wet
scrubber efficiency, all subsequent data were corrected to an
arbitrary reference dust loading of 8 gr/CF. This article
initially suggested the contacting-power concept for estimating
efficiency (at an implied constant inlet dust loading concen-
tration) but the stated important qualification of correcting
to a single reference feed concentration has been de-emphasized
or overlooked in all later work on this concept.
Johnson (1955) tested both a wet cyclone scrubber and
a dynamic wet scrubber (Hydro Volute) and determined that, for
both devices, there was a direct relationship between weight
efficiency and dust loading. For the dynamic scrubber, an in-
crease in fly ash loading from 0.1 to 10 gr/CF air gave an
overall efficiency increase from 94.6% to 98.2%. Similar be-
havior was observed for the cyclone, operated both dry and wet,
and after noting the importance of dust concentration on,the
separation efficiency of inertial collectors, Johnson advanced
the explanation that "inertia1 collectors tend to concentrate
the dust as a first step in the separation process and this
affords an opportunity (by crowding the particles together)
for agg10meration".
Krista1 (1957) tested a Solivore scrubber (low-energy
multi-Venturi unit) and found that for "finer aerosols", such
as CaC03' the increase in collection efficiency with increased
dust loading was significant. As will be shown below, the ap-
F-19
-------
parent size distinction in dust behavior was the result of an
initial dispersal problem in the feed unit, rather than a
real parameter, and with allowance for this dispersion effect,
all efficiency data consistently showed the loading dependency.
The effect of dust loading was explained on the basis of ag-
glomeration: "increased retention with higher dust loadings
is probably due to the greater opportunity for agglomeration
as a result of higher particle concentration".
It is interesting that neither Kristal nor Johnson
examined the fundamental anomaly inherent in their explan-
ations. Assuming the correctness of their statements that
higher inlet dust loadings yield higher agglomeration rates,
why should this yield greater percentage efficiencies based
on gross inlet dust concentrations? If agglomeration rates
are simply proportiona1 to inlet dust concentrations, then
greater percentage removal at higher inlet dust loadings
would require longer scrubber residence times as well, which
is not consistent with the generally constant gas flow con-
ditions employed for these tests. If equipment residence time
remains constant, the only apparent way that percentage ef-
ficiency would increase with inlet gross w~fght dust concen-
tration is for the agglomeration rate to be more than stmply
proportional to dust loadings. In order for a power or ex-
ponential function dependency of agglomeration rate on gross
weight dust loading to prevail in a dispersed-water contactor,
then multi-body collisions (three or more) between dust-dust
and water-dust particles would have to be dominant. Addition-
ally, if this mechanistic explanation is correct, then there
probably should be a minimum dust concentration above which
the efficiency-dust loading dependency would be observed. This
possibility and the nature of the dust-loading effect can best
be ex~lored by examining in detail some of the relevant lit-
erature wo'rk. .
1.
Data of Ingels, Shaffer and Danielson (1960)
Perhaps the most interesting data on the effect of
inlet dust loading and water/gas ratio on scrubber performance
is the commercial test data presented in Chapter 7 of Daniel-
son (1967). These data were taken from the original study' by
Ingels (J960) on multi-stage centrifugal spray scrubbers and
baffled tower spray scrubbers in service on hot-mix as~halt
paving batch plants. Danielson indicates that the significant
dust variable is the minus200-mesh fines feed rate to the
dryer, and it is necessary to refer to the original 1960 study
to determine the derivation of this statement.
F-20
-------
The original article states that the primary dust
v.ariable governing scrubber efficiency was the absolute
scrubber inlet dust loading in lb/hr, with higher efficiency
being obtained at higher abs01ute dust loadings. The scrubbers
were preceded by cyclone pre-cleaners, and it was noted that
"scrubber efficiency was so dependent on the degree of pre-
cleaning that the effect of other variables on collection
efficiency was completely masked ,n the available data".
This effect was graphically shown by Ingels as Figure 4-1,
a plot of scrubber efficiency vs. absolute dust loading.
Absolute dust loadings were used despite the fact that gas
flow rates for the plotted test data ranged from 16,100 to
28,300 SCFM, which would certainly yield different relative
inlet dust loading concentration values. In reviewing these
data, Friedrich (1969) calculated and tabulated the inlet
dust loading concentrations, in gr/CF, utilizing the flow
rates and the absolute emission data, but limited his comment
to noting the wide variations in scrubber inlet loadings and
the absence of information on the design of the primary.col-
lectors operating as pre-cleaners.
Figure 4-1, as presented by Ingels, strongly re-
sembles a standard grade-efficiency curve, with an almost
vertical rise in efficiency from less than 85% at loadings
of 100 lb/hr to over 95%"at 300 lb/hr. Further increases
in loadings from 1000 to 3000 lb/hr show only a gradual ef-
ficiency increase from 97% to 99+%. This plot does not re-
veal any basic efficiency difference due to scrubber type,
or nature of fuel used in the dryer, but the extreme sensi-
tivity of efficiency to dust loading in the range below
400 lb/hr is very apparent. Utilizing the inlet dust load-
ing tabulations of Friedrich (1969) for the Ingels' data,
the data of Figure 4~1 were re-plotted in terms of weight
per cent penetration ~. inlet dust loadings in gr/CF as a
log-log plot in Figure 4-2. Tabulated values for these
runs are also presented in Table 4-1, and the penetration
correlation approach was chosen because of the obvious lack
of efficiency sensitivity in Figure 4-1 at the higher ef-
ficiency levels.
Despite the point scatter in Figure 4-2, which is
not unexpected in view of the extreme variation in the other
parameters (water/gas ratio, gas flow rate, scrubber type,
dryer fuel), the data appear to follow a straight line of slope
= -1. This indicates the penetration to be inversely pro-
portional to inlet dust concentration, which is not at all
evident in Ingels' absolute dust-loading plot, Figure 4-1.
This finding of an inverse proportionality between penetration
F-21
-------
~
..
>-
u
z
w
U
u.
u.
.W
a::
w
co
co
::>
a::
u
en
100
FIGURE 4-1
\VET SCRUBBER EFFICIENCY
ASPHALTIC CONCR,ETE PLANT DUST
DATA OF INGELS, ET AL (1960)
95
90
85
o
80
o
a-CENTRIFUGAL SPRAY
e- BAFFLED TOWER
1,000
2.000 3,000 4,000 5,000
SCRUBBER INLET DUST LOADING, . LBS./HR.
F-22
-------
TABLE 4-1
TEST DATA FROM ASPHALTIC CONCRETE PLANTS
CONTROLLED BY SCRUBBERS
DATA OF INGELS (1960)
Test Inlet Penetration (W /G) (CcJ}(W /G) Fuel
No. Dust % gal/MCF
Loading~
gr. / CF
(a) Multiple Centrifugal-Type Spray Chambers
C - 35 7 4.95 2.2 6.62 32.7 Oil
C-82 2.51 8.4 3.94 9.9 Oil
C - 3 79 18.3 0.9 6.38 116.8 Oil
C-355 9.84 2.2 6.81 67.0 Oil
C - 372B 0.773 15.8 10.99 8.5 Oil .
C-372A 0.492 13.2 11.11 5.46 Gas
C-369 2.32 7.0 5.41 12.55 Oil
. C - 379 15.85 0.8 5.92 93.8 Gas
C-337 . 1.46 4.5 11.11 16.2 Oil
C-426 6.7 1.0 7.75 51.9 Oil
C-417 . 2.56 7.2 2.94 7.53 Oil
C-425 3.14 6.8 4.26 13.4 Oil
C - 385 1.24 8.3 4.56 5.65 Oil
C-433 . 1.58 4.2 8.12 12.8 Gas
------------------------------------------------------------------------
(b) Baffled Towe r Scrubbers
C-393 25.5 0.7 H 12.01 306 Oil
','
C-185 10.2 1.3 19.40 198 Oil
C-234 2.52 5..7 5.70 14..4 Gas
C-418 23.2 0.9 8.90 208 Oil
F-23 .
-------
100
80
60
40
20
10
~ 8
z 6
o
- 4
.-
<.
a:
t-
IJ.J 2
z
IJ.J
Q.
0,4
Q2
. 0.1
0.1
FIGURE 4-2
PENETRATION AS FUNCTION
OF
SCRUBBER INLET DUST LOADING
DATA OF INGELS (1960)
HOT-MIX ASPHALT PAVING PLANTS
~e
o
00
o
o ,
0'0
O-MULTiPLE CENTRIFUGAL SCRUBBER l'LOPE--1
a-BAFFLED-TOWER SCRUBBER
0.3 0.60.7 1.0
6 7 10
30 6070100
3
SCRUBBER INLET DUST LOADING, GRAI~Cr'
F-24
-------
and inlet gross loading is, of course, highly unusual not
only in itself, but also because of the lack of any prior
indication of such a relationship in any series of scrubber
tests or for any piece of scrubber equipment. If it could
be generalized or proven valid, it would have extreme sig-
nificance for correlating wet scrubber performance and ration-
alizing test methods. As will be shown below, re-examination
of the literature data indicate that, when allowance is made
for feed dispersion variables, this relationship does hold
true for most reported wet ~crubber test data.
In this particular case, as is true of most field
test data, inlet dust particle size distribution is also a
variable, and this complicates validation of Figure 4-2 data.
The only information provided by Ingels on this latter vari-
able is ,that of Table 4-2, and these data raise a number of
basic questions. In order to get to the effect of the sec-
ondary variables on efficiency, Ingels avoided the inlet dust
loading "masking" effect with the argument that lithe fractional
collection efficiency of particles larger than 10 microns
(in the scrubber) proved to be 99.7 percent. Consequently,
the variables and operating conditions which affect the amount
and collection efficiency of the 0 to 10 micron fraction
should be reflected in the absolute stack emissions". All
of the data presented in Danielson (1967) on these scrubbers
were derived on this basis. However, Ingels' particle size
analysis of Table 4-2 not only do not support this argument,
but lead to the opposite conclusion. Ignoring the agglomerated
44+ micron sample data, it is clear from Table 4~2 that only
in Test Series C-393 is the collection efficiency of the plus
10-micron dust above the stated 99.7 percent level. In Test
Series C-369, C-372A, C-372B and C-422(1), the 10 to 44-micron
fraction collection efficiencies are considerably less than
99 percent.
With respect to the plus 44-micron fraction, which
shows an anomalously lower fractional efficiency than was
obtained for the smaller dust size fractions, it was estab-
lished by microscopic observations that this fraction con-
sisted of wetted, agglomerated fines. Curiously, the one
set of test data in Table 4-1 that shows no agglomerated
fines in the 44+ micron fraction of the scrubber outlet,
C-393, is for the baffled tower scrubber; all of the other
tests in Tab~e 4-2 are for multiple centrifugal spray chambers.
While this evidence is rather thin, the possibility arises
that the wetted fines and agglomerates are more effectively
removed in a baffled unit than in a centrifugal spray chamber.
F-25
-------
TABLE 4-2
FRACTIONAL COLLECTION EFFICIENCY DATA FOR
SCRUBBERS SERVING ASPHALTIC CONCRETE PLANTS
DATA OF INGELS (1960)
Dust Particle Inlet, Outlet, Efficiency, Inlet, Outlet, Efficiency,
Size, Microns % % % % % %
Test Report Series, C-393 Test Report Series, C-369
.." 0-10 13.0 99.3 95.2 76.4 79.9 92.8
I 10-20 71.1 0.0 100.0 6.3 3.8 96.0
N
0"1 20-44 9.6 0.0 .100.0 2.8 2.0 .95.0
44+ 6,.3 0.7 99.3 14.5 14.3a 93.1
Test Report Series, C-372A Test Report Series, C-372B
0-10 78.0 83.0 85.0 91.0 82.0 85.7
10-20 18.0 5.0 96.2 9.0 3.0 99.4
20-44 2.0 1.0 93.3 0.0 2.0
44+ 2.0 1l.Oa 26.5 0.0 13.0a
Test Report Series, C-422(1)
0-10
10-20
20-44
44+
80.4
18.6
1.0
0.0
73.2
5.1
4.5
17.2
a Microscopic examination indicated that the outlet samples were agglomerated.
-------
ConverselYt Ingels found that the centrifugal spray chamber
was about 5.0 lb/hr more effective than the baffled tower
in' reducing the absolute stack emission. This seeming con-
tradiction may be resolved if it is assumed that the wetting
(agglomerate-formation) and removal functions of a wet
scrubber are consecutive and independent. Thust the data
would indicate that the centrifugal spray units employed
are more efficient in contacting the dust particles with
water, but less effective than the baffled units in remov-
ing all of the wetted material prior to exhaust. The view
of a dispersed-liquid scrubber as a dual-functional device
involving consecutive stages of agglomerate-formation and
agglomerate-removal, with separate and characteristic stage
efficiencies, may serve to rationally differentiate equipment
performance.
2.
Data of Johnson
et al (1955)
Johnson's data on the penetration-dust loading re-
lationship for both dry and wet runs on a cyclone scrubber
operating on micronized talc (mass median diameter = 1.4
microns) are presented in Figure 4-3. Againt an inverse
relationship between penetration and dust loading is shown
(for both dry and wet operation), but the slope of the line
for wet operation is -0.20, which is considerably less than
the -1.0 indicated by Ingels' industrial data, and shows a
fractional exponential effect of dust loadingt Cd, on pene-
tration. It is important to note that the "Harvard gen-
erator" was used in these runs (First, 1953); the test dust
is injected into the system through a dispersing Venturi
after being aspirated off a turntable. If agglomeration
effects are important, as Figure 4-3 certainly indicates
by the fact that dry penetration is an inverse function of
dust loading, then it would seem to be necessary to either
check out the test dust for complete dispersion, and the
absence of pre-existing agglomerates in the feed, or to en-
sure their removal by pre-cycloning (as per Ingels' data).
An inspection of the test system diagram and sampling methods
shows that no precautions were taken against this eventualitYt
and furthert no particle size distribution data on feed
samples are provided. In view of the data of Figure 4-3,
showing the relative ease of self-agglomeration of the dry
talc in a cyclone, the use of another collection device,
the Venturi, as a dispe~ser in this case does not appear
logicalt and it is highly likely that incomplete dispersiont
or secondary agglomeration, was prevalent.
F-2.7
-------
FIGURE 4-3
PENETRATION DIFFERENCE FOR DRY~WET
CYCLONE OPERATION
AS FUNCTION OF INLET DUST LOADING
DATA OF JOHNSON (l955)
~ DRY(NO SPRAY)
50
z SLOPEa.15
0
-
t-
«
0:: 20
t-
W
Z
UJ
Cl. 10
5
.02 ..05 0.1 0.5 1.0 5 10 20
SCRUBBER INLET DUST LOADING, GRAINS/CF
F-28
-------
The use of a Venturi as a feed dust dispersing de-
vice is not confined to the Johnson work. The same device
was used by Lapp1e and Kamack (1955) and by Krista1 et a1
(1957), and as will be shown below, similar lanoma10us" re-
sults were obtained by these investigators for those dusts
which may have had a self-agglomeration tendency. If non-
dispersal or pre-agglomeration of feed dust occurs upstream
of the contacting device on test, then one result would be a
flattening of the penetration vs. dust concentration line,
and the comparative experimental data are consistent with
this possibility. The use of a wet-collector device as a
dry feed dust disperser should properly be viewed with
suspicion in the absence of verifying size distribution data.
In noting that the sl.ope of the dry-operation line
on Figure 4-3 was -0.15, as compared to -0.20 for the wet-
operation line, Johnson assumed that the effect for the dry
line was dust-c;fust agglomeration, and concluded that "some
other effect than agglomeration of dust is contributing to
increased efficiency when water is sprayed. The greater
rise in efficiency with loading (under wet conditions) is a
result of an increase in the relative number of collisions
between particles and drop1ets". It is apparent that "ag-
glomeration" is being implicitly defined solely as dust-
dust collisions, and that water-dust collisions give rise
to "some other effect". There would appear to be a needless
semantic distinction here between the types of multi-body
collisions leading to particle growth. If a certain particle
mass is required prior to wall-collection by the inertial
cyclone forces, then the degree of removal for any given
mode of operation will not be a function of whether this
critical size is attained by dust-dust self-agglomeration
or dust-water inter-agglomeration.
Some additional insight into the probable mech-
anisms at work in Johnson's exper~menta1 work can be ob-
tained by plotting the difference in wet and dry penetration,
bF, of Figure 4-3 against dust concentration, as in Figure
4-4. The slope of the bF vs. Cd line in Figure 4-4 is -0.10,
indicating a slight decrease in the relative effectiveness
of wet operation from bF = 37% at 0.01 grlCF to bF = 18%
at 10 gr/CF. This would be expected on the basis that se1f-
agglomeration of the dust is already at an appreciable
level at high dust concentrations and the addition of more
particles (droplets) would not have as large a relative ef-
fect as when dry particle concentration is too low to yield
appreciable self-agglomeration. However, the relatively
small 50% decrease in bF over a 1000-fo1d range of loading
F-29
-------
FIGURE 4-4
. .
PENETRATION DIFFERENCE FOR DRY.¥! WET
CYCLONE OPERATION
AS.FUNCTION OF INLET DUST LOADING
DATA OF JOHNSON (1955)
100
I
I
-
>-
cr
0
I
....
UJ
3 10 I- -
-
UJ
U
Z
UJ 4 -
cr -
lJJ
I.L.
IJ..
- 2
0 -
z
Q
!;( 1.0 - -
cr
....
UJ
Z
UJ
a..
..
lJ..
<3
0.1
0.01
I
0.10
I
LO
10
SCRUBBi;:R INLET DUST LOADING GRAINS/CF
F-30
-------
values from 0.01 to 10 gr/CF in Figure 4-4 indicates a
probably dominant effect of water droplet concentration
in the wet agglomeration runs. The use of 400 psi hydraulic
nozzles at a flow rate of 6 gal/MCF air probably accounts
for this effect, and this raises the question as to whether
the effect of the dust loading concentration variable should
be considered alone or in conjunction with the water droplet
concentration. If the turbulent agglomeration-collision
mechanism is the means by which the dust particles grow to
a critttal~collection size, then it would appear that col-
lection efficiency under a given force "field would be determined
by the product of the dust and water particle concentrations.
Before examining the data on this basis, it is worth briefly
noting some of the additional literature evidence on the ef-
fect of the water droplet concentration on collection ef-
fi ci ency.
C.
WATER/GAS RATIO
A-correlation approach using water droplet concen-
tration as a prime variable has firm precedence in both theory
and data in wet scrubber performance anal~sis. For example,
Kristal (1957), in reporting the results for the number of
spray stages in multi-stage low-energy Venturi scrubbers,
found that particulate-removal efficiency was dependent on
the spray rate, and concluded that the concentration of water
drops (in the contact zone) was an important factor. In
fact, Kristal stated that "all tests showed an inverse re-
lationship between per cent passage and water rate within the
range tested". As an incidental observation, it should be
noted that this relationship is not consistent with impaction
theory which calls for an exponential (semi-log straight line
plot) relationship between penetration and (W/G) ratios.
Ingels (1960) presented data on the effect of water/
gas ratio for multiple centrifugal and baffled tower scrubbers,
showing that increasing (W/G) reduced the absolute stack emis-
sions. It was observed that low (W/G) ratios were more than
proportionately less effective than higher ratios. This non-
linearity of the (W/G) effect on efficiency was not investi-
gated further, although the spacing of the constant (W/G)
lines 6n Ingels' efficiency plot suggest a logarithmic pro-
gression from low to high (W/G). The use of (W/G) ratios,
usually expressed as gallons/1000 cubic feet of gas flow, is
generally assumed to be a relative measure of the water drop-
let concentration for a given test unit, although this is not
strictly true for all water injection means. However, for
most injection systems which depend on gas turbulence-induced
liquid dispersion, or where the (W/G) ratio varies by reason
F- 31
-------
of the change in gas flow at constant liquid injection rate,
as in Ingels' data, the droplet concentration should be directly
proportional to (W/G).
Lapple and Kamack (1955) presented data on pres-
sure drop as a function of (W/G) for several test scrubbers,
but peculiarly, did n'ot .make similar presentations for dust
penetration as a function of this variable, although such
determinations were made. Instead, penetrations were plotted
as functions of inlet dust loading, with (W/G), as a secondary
"floading" parameter. Inspection of these latter plots show
that, at constant dust loading, penetration decreased 'with in-
creasing (W/G), although the data are too limited to attempt
cross-plots.
In Johnson's data (1955), the maximum efficiency levels
were obtained in a wet cyclone scrubber with the highest spray
nozzle atomization pressure of 400 psig, and the efficiency
levels at lower pressures were observed to be roughly propor-
tional to the nozzle pressure. Also, for a given nozzle pres-
sure, the collection efficiency was a linear function of cyclone
inlet gas velocity, except for the extremely high 400 psig noz-
zle pressure run. These data indicate both the effect of water
droplet concentration and turbulent gas mixing on collection
efficiency. Unfortunately, the pre-agglomeration of the feed
dust compromis~s the results, and it is questionable'ff the
data can be use~ for more than qualitative behavior comparison.
Review and re-plotting of the data covered by the
above investigations definitely established that the combined
dust and water particle concentration terms, expressed as a
product, (Cd)(W/G), served to correlate the effect of these
two primary variables on the efficiency. This correlating
parameter has been termed the "agglomeration index", and its
utility is explored below.
D.
AGGLOMERATION INDEX
Particle growth to a critical minimum size neces-
sary for removal from the gas phase by inertial wall collection
or other mechanisms may be considered to occur primarily by
dust-water collisions, and the two usual test variables, inlet
dust loading and water/gas ratio, can be taken as (approximate)
particle concentration terms. A theoretically-correct ex-
pression for collision probability would involve the product:
(No. dus~ particles/CF) (No. water drops/CF). Expressing
dust loadings and water/gas rates as number concentrations
would require correcting the mass or volumetric concentrations
F-32
-------
for the cube of the mean particle diameter, the further cor-
rection of the dust loading for solids density. However, for
a given test dust and water-injection system, the usual re-
spective (Cd') and (W/G) units of grains/CF and gallons/MCF
may be used directly in the form of an "agglomeration index",
Xc' such th at:
wh e re
Xc = (Cd)(W/G)
Cd = dust concentration/unit gas
volume, commonly gr/CF
WIG = water/gas volumetric ratio,
commonly gal/MCF
( 4- 1 )
Underlying the postulation of Equation (4-1) are
several general assumptions. It is assumed that the primary
water-dust contacting mechanism is turbulent mixing, in agree-
ment with the work of Semrau (1958), and that such contact-
ing will be a statistically random process. It is further
assumed that the water drops are larger than the dust particles
and above the critical collection size, while the dust will
be below this size. This will be true in most wet scrubbing
units, and therefore the dust-dust collision processes may be
ignored. However, as already seen, if critical-size dust
pre-agglomerates exist in the feed, different behavior may
be expected than for well-dispersed dusts. Because the pri-
mary mixing mechanism is turbulence, the agglomeration index
can only be used as a correlating parameter at constant
levels of turbulence scale and intensity. For a given
scrubbing unit, this implies the proper application of the
Xc parameter at constant gas velocity and/or pressure drop.
Another reason for specification of constant gas velocity
for use of the Xc parameter is the interaction between gas
velocity and water drop stability. Above a certain range of
air velocities, roughly 200 ft/sec, drop-shattering pre-
dominates, and drop size is a function of the gas velocity.
In this context, Equation (4-1) is only a partial index of
collision probability, and a gas turbulence or velocity
function term would be required for complete description of
the correlating parameter. This final variable will be
covered in a later section of this report. However, in
order to test Xc within the limitations outlined, it is in-
formative to return to the data of Ingels (1960) to test the
applicability of this approach.
F-33
-------
1.
Data of Ingels (1960)
Values of (Cd)(W/G) were calculated for Ingels' in-
dustrial test data, and are included in Table 4-1. When the
table 4-1 data are now plotted as per cent penetration vs.
the agglomeration index, Xc' as in Figures 4-5, 4-6 and-r-7,
separation of the test data on the basis of dryer fuel and
scrubber type is obtained. While Figure 4-5 still exhiBits
point scatter, the best line through the data for the multi-
stage centrifugal scrubbers (oil-fired dryer) gives the same
-1.0 slope shown in Figure 4-2, where penetration was plotted
against inlet dust concentration alone. Figure 4-6, a similar
plot for the baffled-tower, gas-fired scrubbers, shows that
these points fall parallel to, but lower than, the line for
the centrifugal scrubbers, oil-fired. The higher penetration
shown for the baffled-tower scrubbers compared to the centri-
fugal spray scrubbers isin accordance with the findings of
Ingels (on the basis of absolute stack emissions and a curvi-
linear multiple correlation) that the latter type of scrubber
is the more efficient in this instance.
Figure 4-7 compares the gas-fired dryer data for
the two types of scrubbers with the lines for the oil-fired
units, established in Figures 4-5 and 4-6. Although the data
are quite limited, penetration for the gas points appear to
be less than for oil-firing, especially for the baffled-tower
equipment (one point). Again, this is in' agreement with the
conclusions of Ingels, based on his entirely different cor-
relation apprDach~ that stack emissions were higher when the
dryer was oil-fired, probably due to additional particulate
matter contributed by the fuel oil. It should be noted that,
while point scatter may impair the validity of the conclusions
draw~ from the plotted data, Ingels also stated that the. data
scattered badly, "even when corrected for the variables studied".
The fact that the use of the (Cd)(W/G) correlation
group allows separation of the fuel and scrubber type vari~
ables in Ingels' data is an indication of the potential
utility of this approach. Essentially, the same set of con-
clusions may be derived from the simple penetration vs. X
plots as was obtained by Ingels by statistical analysis o~
the same experimental data. Further generalization of this
correlation technique requires additional data review, and
a comprehensive series of wet scrubber test data has recently
been provided by Calvert and Legatski (1970). These data
allow a remarkable demonstration of the utility of the ag-
glomeration index approach to comparative scrubber evaluation.
F-34
-------
FIGURE 4-5
PENETRATION AS FUNCTION OF
AGGLOMERATION INDEX, (Cd) (WIG)
DATA OF INGELS (1960)
MULTIPLE CENTRIFUGAL SCRUBBERS
OIL- FIRED DRYER
100
o
10
~
z
0
~
I:r
...
W 0
Z
W
Q.
1.0 0
0.1
1.0
10
100
'Xc,AGGLOMERATION INDEX -(Cd)(W/G).
F-35
1000
-------
100
~
z
o
....
«
~
UJ
Z
UJ
n.
FIGURE 4-6
PENETRATION AS FUNCTION OF
AGGLOMERATION PARAMETER,(Cd)(W/G)
DATA OF INGELS ( 1960)
. .
MULTI P LE CENTRI FUGAL SCRUBBERS
GAS-FI RED DRYER
10 -
1.0 f-
QI
1.0
I
I
,
,
,
0'
"
,
"
,
,
,
,
o , "
,
" rL1NE FOR OIL-FIRED
'V DRYER
,
,
"
"
,
,
0',
,
,
O-GAS-FIRED DRYER
I
10
,
100
Xc ,AGGLOMERATION INDEX- (Cd) (WiG)
F-36
-
-
1,000
-------
FIGURE 4-7
. .
PENETRATION AS FUNCTION OF
AGGLOMERATION INDEX, (Cd) (WIG)
DATA OF INGELS ( 1960)
BAFFLED TOWER SCRUBBERS
100
I
I
10 I-
,
,
,
,
,
,
,
"
,
'~
0,
~
z
o
-
I-
~
I-
W
Z
W
Q.
e
1.0 ~
0- OIL- FIRED DRYER
a-GAS-FIRED DRYER
0.1
0.1 .
I
I
100
1.0
10
XCI AGGLOMERATION INDEX-(C'd)(W/G)
F- 37
-
-
-------
2.
Data of Calvert and Legatski (1970)
This investigation covered the experimental per-
formance, in terms of penetration, of a series of wet scrubbers
tested against relatively low concentrations of respirable
coal dust. The scrubbers tested included:
(l) Venturi
(2) IIVentri'~Rodll.
(3) IIVentri-Spherell
(4) Impingement Plate
( 5) IIAir Tumblerll
(6) Packed Bed
(7) Multiclone
( 8) IIWetted Screenll
Coal dust concentrations fed to the test scrubber were of
the order of 0.01 grISCF. This relatively low level together
with pre-removal of the oversize by use of an upstream cyclone
in series with the scrubber ensured the absence of agglomerates
in the scrubber feed dust. The particle size of the test coal
dust was about 1.5-2.5 microns, with standard deviations of
the same order of magnitude.
2 ( a ) .
Venturi Scrubbers:
and IIVentri-Spherell
Venturi, IIVentri-Rodll
The Calvert and Legatski data afford a rigorous
test of the proposed (Cd}(W/G) correlation parameter, and the
calculated Xc values for the Venturi and IIVentri-Rodll Scrubber
test data are presented in Tables 4-3 and 4-4, and Figure 4-8.
Comparison of Table 4-3 and 4-4 with Figure 4-8 will show that
the points plotted are the individual samples composing a
single run. In view of the fact that the Venturi data were
obtained at only one (WIG) level, and the IIVentri-Rodll data
at only two, the only factor providing an adequate Xc range
in Figure 4-8 was the random variation in inlet dust loading.
As Figure 4-8 indicates, the gross penetration values associ-
ated with the random variations of the independent variable
are quite clearly dependent on these variations. It is rather
F-38
-------
TABLE 4-3
GROSS PENETRATION FOR A VENWRI SCRUBBER
FOR AN INLET DUST OF MEAN DIAMETER 1.35 MICRONS
AND STANDARD DEVIATION 1.7
DATA OF CALVERT AND LEGATSKI (1970)
Concentration
Run Configuration Water Rate Sample (C(j){W /G)(l) Penetration
gal./MCF No. (mg./SCM) (gr. /SCF) %
." 79 No Rods 500 And. (2) 40.9 0.0179 0.089 42
I
w 435 52.8 0.0231 0.115 39
\0
438 48.8 000213 00106 39
80 No Rods 5.0 441 48.4 000212 0.106 36
And. 38.1 0.0166 0.083 43
444 4808 0.0170 0.085 37
446 5205 000229 00114 35
Average 4702 0.0206 00103 38
81 No Rods 10.0 And. 46.3 0.0202 0.202 21
449 50.2 ,0.0219 00219 21
452 5109 0.0227 00227 19
454 5207 000230 0.230 17
Average 5003 0.0220 19
-------
TABLE 4-3 Cont'd.
Concentration
Run Configuration Water Rate Sample (Cd)(W /G) Penetration
gal. /MCF No. (mg. /SCM) (gr. /SCF) %
82 Rods 5.0 459 47.4 0.0207 0.104 44
And. 37.5 0.0164 0.082 49
462 44.9 0.0196 0.098 46
464 47.3 0.0207 0.104 41
83 Rods 5.0 And. 33.8 0.0148 0.074 48
"T'\ 467 42.4 0.0185 0.092 42
. 470 43.1 0.0188 0.094
~ 40
o 472 43.0 0.0188 0.094 40
Average 42.4 0.0185 0.092 43
(1) (gr 0 /SCF) {gal o/MCF)
(2) Andersen Sampler
-------
TABLE 4-4
GROSS PENETRATION FOR A "VENTRI - ROD" SCRUBBER
FOR A COAL DUST OF MEAN DIAMETER 1.42 MICRONS
AND STANDARD DEVIATION 1.64
DATAOF CALVERT AND LEGATSKI (1970)
Concentration
Run Water Rate Sample No. (Cc1>(W /G)(1) Penetration
." sa1./MCF (mg./SCM) (gr .. /SCF) %
I
~ 86 6 489 31.2 0.0136 0.0817 8~3
-
And. (2) 23.3 0.0102 0.0612 11~6
492 28.3 0.0124 0.0744 9.9
494 31~0 0.0136 0.0816 8.4
Average 28.5 0.0125 9.4
87 10 Ando 19.5 0.00852 00085 6~4
497 26.~ 0.0115 0..115 5.8
500 31.9 0..0139 00139 5..5
88 10 503 34.0 0.0149 00149 5..2
Ando 25.6 0..0112 0.112 5..9
506 34.3 0.. 0150 0..150 5..2
89 10 And. 23.4 0.0102 00102 701
509 32.2 0.0141 00141 5..6
512 3500 000153 00153 500
514 3400 000149 0..149 4.7
Average 29.6 0.0129 5.5
(1) (gr./SCF){gal./MCF)
(2) Andersen sampler
-------
o~
z
o
....
o-1
"VENTR I-ROD"
"
,
~
~Sl.OPE..1
THROAT
VELOCITY
FPS
VENTURI 223
"VENTRI-ROD~3 70
+ RODS INSERTED
0.02
0.04 0.06 0.1 0
0.08
0.4 Q6 o.a 1.0
0.2
4CI AGGLOMERATION INDEX. (Cd) (WIG)
F-42
-------
remarkable that the agglomeration index approach is sensi-
tive enough to show correlation for what would ordinarily
be presumed to be random experimental sample error, as was
assumed by Calvert and Legatski.
Despite the absence of deliberate changes in the
primary variables, and the necessary use of supposedly in-
cidental variations, both sets of data correlate quite well
with the -1 slope found for the Ingels' asphalt plant scrubber
data, substantiating the earlier tentative conclusion of a
simple inverse relationship between gross penetration and
Xc. The relationships shown in FigOre 4-8 are of the form:
F =K
(Cd)(W/G)
(4-2)
where
F = penetration, weight percent
K = proportionality constant
At constant gas velocity or pressure drop, Equation
(4-2) can obviously be used to compare scrubber performance.
The "K" value may be looked upon as a relative performance
index, with a smaller value denoting superiOr performance.
For the data of Figure 4-8, using the conventional units for
Cd and (WIG):
Venturi:
K = 4.13
K=0.724
"Ventri-Rod":
This K comparison, indicating lower penetration values for
the "Ventri-Rod" unit as compared to the Venturi, is illus-
trative and not absolute because of a substantial velocity
difference in the two sets of runs. Part of the performance
difference indicated, by the relative K values may be at-
tributed to the fact that the "Ventri-Rod" scrubber was op-
erated at a throat velocity of 370 fps, which is 70% greater
than the 223 fps velocity used in the Venturi. However,
even after correcting for the velocity differential on the
basis of the velocity ratio squared, the "Ventri-Rod" unit
still shows substantially better performance than the Venturi
using the agglomeration index approach. It is interesting
that no comparable conclusion was drawn by Calvert and
Legatski, using the conventional impaction theory approach.
F-43
-------
In contrast to the direct tests on the commercial
"Ventri-Rod" unit some additional limited data were run on
a home-made Venturi, with and without throat rods. In the
test series on the Venturi scrubber, Table 4-3, two runs
(eight samples) were made with rods inserted below the throat,
and Calvert and Legatski concluded that the presence of these
rods had a negative effect on the scrubbing efficiency.
Figure 4-8 shows that, as a group, the penetrations for the
rods-inserted runs were slightly greater than for the Venturi
runs made without the rods. However, it is quite clear that
the reason for this apparent shift is the lower values of
the agglomeration index(run conditions) and that the rod data
correlate with the rest of the data, and their insertion is
without effect on penetration. The agglomeration index would
thus seem to have a good deal of potential as an analytical
tool at the experimental level in differentiating between
random or real effects. Further, it is apparent that a valid
performance evaluation requires comparison under similar Xc
conditions.
Another modification of the Venturi that was tested
by Calvert and Legatski was a "Ventri-Sphere" scrubber, which
consists of a Venturi section with a long diffuser section, a
1800 gas reversal, and a l-foot deep flooded bed of plastic
spheres in the annular space surrounding the ,Venturi. -Data
for these test~ are provided in Table 4-5 and-are plotted in
Figure 4-9. The tests were made at two different Venturi
throat velocities, and despite the absolute values of pene-
tration, the data appear to confirm the expectation of a lower
penetration line for equal values of Xc at the higher throat
velocity. However, point scatter is severe at these lower
penetration levels (higher Xc) and it is apparent that the
combination of inlet dust concentrations of the order of 0.02
gr/CF and removals of 99% have reached the limits of experi-
mental measurement capability. Despite this, both lines in
Figure 4-9 were drawn with a -1 slope and appea~ to reasonably
represent the data. The penetration data for the "Ventri-
Sphere" fall below the line in Figure 4-8 for the "Ventri-
Rod" scrubber, indicating that, at projected similar throat
velocities, the flooded bed does actually contribute some
additional dust removal capability.
2(b).
Impingement Scrubber
The impingement scrubber tested on the coal dust
consisted of two stages of standard-design impingement plates,
with a tangential gas inlet (cyclone effect) below the stages.
F-44
-------
TABLE 4-5
GROSS PENETRATION FOR A "VENTRI-SPHERE" SCRUBBER
FOR AN INLET DUST OF MEAN DIAMETER 1.44 MICRONS
AND STANDARD DEVIATION 1.7
DATA OF CALVERT AND LEGATSKI (1970)
Concentration
Run Water Rate Air Rate Sample (C d) (W /G)(l) Penetration
gal. /MCF CFM No. (mg. /SCM) (gr. /SCF) %
.."
I
~ 95 16.7 1500 And. (2) 48.2 0.0210 0.350 1.0
U'1
547 58.7 0.0256 0.427 0.6
550 63.0 0.0275 0.459 1.4
96 16.7 1500 553 59.0 0.0258 0.431 1.4
And. 52.0 0.0227 0.379 1.5
556 64.3 0.0281 00469 1.1
97 16.7 1500 And. 52.6 0.0230 0.384 2.2
559 65.9 0.0288 0.481 1.1
562 67.6 0.0295 0.493 100
Average 59.0 0.0258 0.431 182
98 5.0 2000 565 53.1 0.0232 00116 2.9
And. 4605 0.0203 00101 3.3
568 56.1 000245 0.122 2..6
-------
."
I
.::-
0'\
TABLE 4-5 Cont'd.
Concentration
Run Water Rate Air Rate Sample (CcJ)(W /G) Penetration
ga1./MCF CFM No. (mg. /SCM) (gr. /SCF) %
99 5.0 2000 And. 41.0 0.0179 0.089 4.3
571 55.2 0.0241 0.120 2.0
574 56.7 0.0248 0.124 2.6
Average 51.4 0.0224 0.112 2.9
(1) (gr. /SCF)(gal. /MCF)
(2) Andersen Sampler
-------
FIGURE 4-9
IIVENTRI- SPHERE" SCRUBBER
GROSS PENETRATION AS FUNCTION
OF
AGGLOMERATION INDEX, (Cd) (WIG)
-DATA OF CALVERT a LEGATSKI (1970)
10.0
8
6
4 "Z
~ ~ e
z 2 ,
o '~ea
-
I-
0.4
VELOCITY
0-414fps
e- 276 tps
.02
0.1
.
0.01
002
0.04 0.06 0.10
0.08
0.2
0.4 0.6 0.8 1.0
XC, AGGLOMERATION INDEX -(Cd)(W/G)
F-47
-------
The scrubber was 2.5 ft in diameter, and the trays had 22%
open area, yielding a free flow area of 0.68 ft2. The unit
was operated with a countercurrent liquid flow rate (sprayed
in) of 20 GPM and in one set of runs this was augmented with
a 20 GPM cocurrent flushing spray directed at the bottom
plate. The data for these runs are presented in Table 4-6
and Figure 4-10.
On an impingement plate, the atomization of the
water occurs at the tray orifices, and the degree of this
atomization is nominally a function of the liquid depth, and
hence the counter-current liquid rate. The bottom spray
serves to keep the underside of the bottom plate washed free
of accumulated solids, and since it contacts the gas at low
gas energy levels, the contribution of this flushing liquid
rate would not normally be counted toward (~/G). This pro-
cedure was originally followed in plotting the data of Table
4-6. However, it was. found that when all of the data of
Table 4-6 were plotted in this manner, without differentiation
for the use of the auxiliary spray, it was virtually impos-
sible to discern any pattern in the point grouping. There-
fore,"the data were separated into two groups, depending on
whether or not the bottom spray was used, and were re-plotted
as shown in Figure 4-10.
Figure 4-10 shows that the set of penetration data
obtained without the bottom spray conforms to the reciprocal
Xc relationship already established, and further, that the
comparative penetrations for the two-spray data are all lower
than the best penetration-Xc line for the single-spray data.
Thus, it may be concluded that the bottom spray is providing
a small, but definite, incremental contribution to the effic-
iency of the unit. Th~s is in direct contradittion to the
conclusions of Calvert and Legatski, who, on the basis of the
original Table 4-6 data, decided that lithe gross penetration...
appears to be essentially unaffected by the plate spray for
this particular dust". Again, this illustrates the utility
and sensitivity of the agglomeration index approach, and its
general applicability regardless of equipment type. The
relatively poor performance of the impingement scrubber as
compared to the various Venturi units discussed above may be
attributed to the fairly low maximum (orifice) velocity of
61.2 fps.
2(c).
"Air Tumbler"
This device is a horizontal wet cyclone contain-
ing a water well, and the cyclonic action of the air causes
F-48
-------
TABLE 4-6
GROSS PENETRATION FOR AN IMPINGEMENT SCRUBBER
FOR AN INLET DUST OF MEAN DIAMETER 1.44 MICRONS
AND STANDARD DEVIATION 1.7
DATA OF CALVERT AND LEGATSKI (1970)
Run Spray Water Rate Sample Concentration (CcJ)(W /Gj) Penetration
gal./MCF No. %
Top Bottom (mg./SCM) (gr. /SCF)
90 8 8 517 35.0 0.0153 0.122 42.3
.." And. (2) 29.2 0.0128 0.1024 43.8
I
~ 520 36.2 0.0158 0.126 40.3
\0
91 8 8 And. 25.4 0.0111 0.0888 46.1
523 31.6 0.0138 0.1104 39.6
526 33.5 0.0146 0.117 38.2
92 8 8 529 33.5 0.0146 0.117 40.9
And. 28.6 0.0125 0.100 42.7
532 34.9 0.01525 0.122 40.1
Average 32.0 41.4
93 8 0 And. 31.4 0.0137 0.110 52.2
535 36.4 0.0159 0.127 40.1
538 39.0 0.0170 0.136 4308
-------
.."
I
c.n
o
Run
Spray Water Rate
gal./MCF
Top
Bottom
94
8
o
(I) (gr./SCF}(gal./MCF)
(2) Andersen Sampler
Sample
No.
531
And.
544
Average
TABLE 4-6 Cont'd.
Concentration
(mg./SCM)
36.4
34.2
37.1
(gr. /SCF)
0.0159
0,,0150
0.0162
35.8
(C&(W/G)
0.127
0.120
0.130
Penetration
%
42.3
46.2
41.5
44.1
-------
FIGURE 4-10
2-STAGE IMPINGEMENT SCRUBBER
GROSS PENETRATION AS FUNCTION
OF
AGGLOMERA TION INDEX, (Cd) (WIG)
DATA OF CALVERT a LEGATSKJ (1970)
100
80 BOTTOM FLUSH SPRAY USED
60 '\. J
" .
" !
40 0c9~
-------
pickup and dispersion of the liquid from the well. In this
mode of operation, the liquid feed rate to the equipment is
not directly related to the actual dispersed-water/gas ratio
in the contact zone. The data for the single run made by
Calvert and Legatski for this unmodified unit are given in
Table 4-7, and the gross penetration of the 1.35-micron dust
was a high 58%. The unit was modified by the inserti6nof a
IIflushing devicell at the gas entrance which raised the point
velocity at this section to 100 fps; this was termed IICon-
figuration 211. A second modification, IIConfiguration 311,
consisted of the addition of a low-energy Venturi section
(throat velocity = 200 fps) between the flushing device and
the IIAir Tumb1erll proper, and data for these runs are also
listed in Table 4-7.
No description of the flushing device or the al-
location of water to the various components in series was
provided, so that it was not possible to calculate any mean-
ingful (W/G) ratios for these runs. Instead, the gross pen-
etration data were plotted against inlet dust loadings in
Figure 4-11, and by separation of the runs by configuration
and gross wate~ rates, a degree of data order is achieved.
For Configuration 2, at "2.5 GPM, the data agree quite well
with the expecta:~ion of a reciprocal penetration-Cd relation-
ship. Raising the gross flow rate to 7.5 GPM (w1th unspeci-
fied distribution) at this same configuration gave a group
of data points showing somewhat higher penetrations, rather
than the lower values to be expected. : Without a description
of the IIf1ushing devicell, it"is not possible to further eval-
uate these negative results. However, for the Venturi
section-addition of Configuration 3, lower penetrations are
obtained at the higher gross water rates, as would be ex-
pected from previous agglomeration index behavior patterns.
In addition to the absence of required information on the
physical aspects of the equipment, no cross-sectional area
data was provided for the main cyclone body, and linear
velocities could not be estimated from the stated volumetric
flow rate of 2200 CFM.
2(d}.
Packed Bed Scrubber and Multic10ne
As part of the scrubber study conducted by Calvert
and Legatski, tests were made on a Multiclone unit, operated
dry~ and a 3-foot deep bed packed with 1-1/211 polypropylene
pall rings, operated both dry and wet. The data for these
runs are tabulated in Tables 4-8, 4-9 and 4-10, and are
plotted in Figure 4-12.
F-52
-------
TABLE 4-7
GROSS PENE1RATION FOR AN "AIR WMBLER"
FOR AN INLET DUST OF MEAN DIAMETER 1.35 MICRONS
AND STANDARD DEVIATION 1.7
DATA OF CALVERT AND LEGATSKI (1970)
Run . Configuration Nominal Sample Concentration (CcI>(W IGr Penetration
Water Rate No. %
gal./MCF ~mg./SCM) (gr. ISCF)
."
I 77 1 7.5 426 38.7 0.0169 0.1268 58
<.11
w
71 2 2.5 374 56.4 0.0246 0.0615 33
And. ** 56.0 0.0244 0.0610 28
377 44.6 0.0195 0.0488 69
379 65.4 0.0285 0.0713 28
72 2 2.5 And. 40.0 0.0175 0.0438 41
382 54.9 0.0240 0.0600 33
385 57.0 0.0249 0.0623 31
73 2 2.5 388 51.4 0.0224 0.0560 36
And. 38.1 0.0166 0.0415 44
391 51.4 0.0224 0.0560 36
393 50.2 0.0219 0.0548 36
Average 51.4 0.0224 0..0560 37
-------
TABLE 4-7 Cont'd.
Run Configuration Nominal Sample Concentration (C&(W/G) Penetration
Water Rate No. %
gal./MCF (mg. /SCM) (gr. /SCF)
74 3 2.5 And. 40.4 0.0176 0.0440 51
396 54.5 0.0238 0.0595 42
399 55.1 0.0241 0.0603 41
401 56.9 0.0248 0.0620 41
"
I
U1 Average 51.7 0.0226 43
~
75 3 10.3 404 66.2 0.0289 0.298 31
407 68.7 0.0300 0.309 30
409 69.1 0.0303 0.312 30
And. 60.4 0.0264 0.272 30
76 3 10.3 413 61.2 0.0267 0.275 33
And. 48.4 0.0211 0.217 37
416 6S~8 0.0287 0.295 32
418 64.2 0.0280 0.288 32
Average 63.0 0.0275 0.283 32
-------
TABLE 4-7 Cont'd.
Run Configuration Nominal Sample Concentration (CcV(W /G) Penetration
Water Rate No. %
gal. /MCF ~mg./SCM) .(gr. /SCF)
77 2 7.5 And. 43.2 0.0189 0.142 31
421 49.5 0.0216 0.162 38
424 48.3 0.0211 0.158 39
" 78 2 7.5 429 68..7 0.0300 0..225 32
. And.. 57.5 0.0251 0..188 33
U1
U1 432 77.8 0.0340 0..255 30
Average 57.5 0..0251 0..188 34
. (gr./SCF)(gal./MCF)
.. Andersen Sampler
1
2
3
Air Tumbler Alone
Air Tumbler with "Flushing Device"
Air Tumbler with Venturi and "Flushing Device"
-------
.~
z
o
~
~
a::
t-
UJ
. Z
W
Q.
(/)
(/)
o
a::
.0
100
80 CONFIGURATION 2 (2.5GPM) .
60
10
100
80
60
40
30
20
40
20
10
0.01
FIGURE 4-11
"AIR TUMBLER"
GROSS PENETRATION AS FUNCTION
OF
INLET DUST LOADING
DATA OF CALVERT 8 LEGATSKI (.,970)
CONFIGURATION 3
.~
"
~
,
0-2.5 GPM
+-IO.3~.pM
CONF}GURATION 2 (7.5GPM)
,
~'O
o ,00rLlNEFOR
',< 2.5 G pM'
002 0.030.04 a06 0.1 o.or
008
0.04 0.06 0.1 .
008
0.02
. ,
INLET OUST lOADING, G.RAINS/SC,F
F-56
-------
TABLE 4-8
GROSS PENETRATION FOR A DRY PACKED BED SCRUBBER
fOR AN INLET DUST OF MEAN DIAMETER 1082 MICRONS
AND STANDARD DEVIATION 1.68
DATA OF CALVERT AND LEGATSKI (1970)
Concentration
Run Sample . Penetration
No. (mg. /SCM) (gr. /SCF) %
"T1
. 57 And.* 11.7 0.00511 40
U1
" 269 10.7 0.00467 30
271 12.0 0.00524 38
58 276 10.2 0.00446 46
278 13.8 0.00603 35
And. 8.8 0.00385 42
Average 11.2 0.00489 39
. Andersen Sampler
-------
Run Sample
." No.
I
U1 63 314
co
316
And. (2)
64 And.
326
328
65 332
And.
334
66 340
And.
.342
Average
(1) (gr. ISCF)(gal. IMCF)
(2) Andersen Sampler
TABLE 4-9
GROSS PENETRATION FOR A WET PACKED BED SCRUBBER
FOR AN INLET DUST OF MEAN DIAMETER 1.44 MICRONS
. AND STANDARD DEVIATION 1.74
WIG = 2.0 gal./MCF
DATA OF CALVERT AND LEGATSKI (1970)
Concentration
(Cd>(W IG)(1) Penetration
(mg./SCM) (gr. ISCF) %
13.6 0.00594 0.0119 29
13.9 0.00612 0.0122 43
12.9 0.00564 0.0113 26
15.4 0.00673 000135 36
15.1 0.00660 0.0132 26
15.7 0.00686 0.0137 36
13..8 0000603 0.0121 27
13.2 0.00577 0.0115 27
15.3 0.00669 0.0134 31
14.4 0.00629 0.0126 .33
10.9 0.00476 0.00952 47
15.6 0.00682 0.0136 38
14015 33
-------
TABLE 4-10
GROSS PENETRATION FOR A MULTICLONE
FOR AN INLET DUST OF MEAN DIAMETER 1.5 MICRONS
AND STANDARD DEVIATION 1.7
DATA OF CALVERT .AND LEGATSKI (1970)
Concentration
Run Sample Penetration
No. (mg. /SCM) (gr. /SCF) %
"
I 67 And. * 12.6 0.0550 82.5
c.n
\0 348 15.3 0.0668 73.9
346 15.5 0.0677 70.3
351 14.1 0.0616 72.3
68 355 14.9 0.0650 77.2
And. 11.6 0.0507 89.7
357 16.4 0.0716 73.2
69** 393 44.3 0.1934 73.8
And. 29.4 0.1284 66.7
360 33.7 0.1472 75.4
365 43.7 0.1909 80.1
70 368 54.6 0.238 60.6
And. 39.7 0.1733 86.1
371 56.1 0.245 80.4
* Andersen Sampler
** A new dust was used starting with Run #69
-------
FIGURE 4-12
DRY 8 WET PACKED BED SCRUBBER
a MULTICLONE
GROSS PENETRATION AS FUNCTION OF
AGGLOMERATION INDEX OR DUST LOADING
DATA OF CALVERT a lEGATSKI (1970)
100
80 PACKED BED
60
~ .
..
40 t
..
6
~.
~ 20 +-DRY
z
0 A-(W/G)a2GAL/MCF
~
~
~ 10
~ 160 ~'"'
w 80 MULTICLONE e'e e
a. e e-,
(f) e ,
60 e,
CI)
0
cr
" 40
o-DUST-I
a-DUS T-2
20
10
0.001
0.002
0004 0.006 0.01
0.008
0.02
0.04 0.06 0.1
0.08
Cd INLET DUST LOADING, GRAINS/SCF
F-60
-------
Both types of scrubbers tested in this series are
inertial impaction units which do not depend on dust-partic1e/
water-droplet collisions for agglomeration to critical re-
moval size. Therefore, the collision function, or agglom-
eration index, use~ for correl~tion of the gross penetration
behavior for drop/particle contacting should not be applicable
to these units. Howev~r, in the absence of dispersed water
drop1ets~ the self-agglomeration of dust in the turbulent or
inertial field should be a function solely of the inlet dust
concentration, Cd, as shown by the data previously plotted.
The results of Figure 4-12 are in accord with this expectation.
The dry packed bed data and the first set of multi clone runs
appear to correlate 'well with the (Cd)-l relation. However,
the second set of dry Mu1tic10ne runs at the higher dust load-
ings shows too great a degree of point scatter to permit more
than an indication of the possible position of the correlating
line. Reference to Table 4-l0,shows that the investigators
made spe'-cific note of the fact that a "new dust" was used for
the second set of Mu1ticlone runs, but gave no reason for
the accompanying shift to higher weight dust loadings. The
relative position of this second set of data on the Cd plot
is evidence that the most probable cause was a marked change
in particle size distribution, but no analytical data on this
point were provided.
The packed bed data of'Figure 4-12 show a fair cor-
relation" on' a Cd basis for the ,dry runs, but a completely random
pattern for the wet runs at a constant (W/G) = 2.5 ga1/MCF.
Because the dry bed causes the gas to follow a tortuous,
curved path, inertial and self-agglomeration behavior similar
to the Multiclone unit would be expected for the dry dust, and
this is substantially the observed behavior. The displace-
ment of the wet-run packed-bed points to the right in Figure
4-12 results from the use of the (Cd)(W/G) index for these
data rather than Cd alone; the latter is the most probably
correct variable in this test sttuation. The wet-run data
scatter may result either from the narrow range of inlet dust
loadings employed, the extremely low values of dust loading,
or from gas backmixing and recirculation caused by the oppos-
ing water flow in the countercurrent contact packed bed.
Within the limitations of the data, there appears to be no
particular advantage, from a gross penetration basis, of
wet countercurrent packed bed operation, as compared to dry
operation at these dust loading levels.
If the scrubber performance is dependent on inlet
dust loadings, as the data have thus far indicated, then it
would appear mandatory that comparative equipment tests be
F-61
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carried out using similar dust concentrations. An inspection
of the tabulated Calvert and LegatsRi data in this report
will show that the packed bed and Multic10ne runs, as well as
those for the "wetted screen" discussed below, were made at
inlet coal dust loadings roughly 25% of those used for the
Venturi and other equipment types. The relatively high gross
percentage pehetration levels of Figure 4-12 associated with
these low loadings are indicative of the seemingly anomalous
data resulting from allowing order-of-magnitude variations in
the inlet dust concentration test conditions. Failure to hold
this primary variable within reasonably narrow test limits
would tend to invalidate any conclusions drawn from data un-
corrected for this variation. Correction to a comoarable Cd
(or X ) basis should be made by the method of Lapple and
Kamac~ (1955) of extrapolating to an arbitrary fixed dust
loading value, or by the dfrect graphical'method employed in
this report.
2(e).
"Wetted Screen" Scrubber
This unit consisted of a pleated, accordion-like
sprayed-screen assembly, followed by a horizontal cyclone
separator. Data for the tests on this unit are presented in
Table 4-11 and in Figure 4-13, and inspection again indicates
that failure to allow for the dust loading dependency ser-
iously compromised the conclusions drawn from the test re-
sults. The data of Figure 4-13 scatter badly, but in view
of the very low X test levels, this is not unexpected.
The low agglomera{ion index values arise not only from the
low dust 10adin9 concentrations of 0.005 gr/SCF, but also
from the small (W/G) ratio of 0.75 gal/MCF. With due con-
sideration for these latter values, the penetration values
obtaineq are remarkably low, indicating excellent unit per-
formance, which is directly contrary to the conclusions of
the investigators. While it is by no means certain from the
daca of Figure 4-13 that the penetration/Xc reciprocal re-
lat10nship applies, if a 45° line drawn through the data is
extended to the higher Xc ranges used for testing the Venturi-
type scrubbers, lower comparative gross penetrations are
inditated for the wetted screen scrubber. However, such
unit comparison on an Xc basis is not fully warranted on
the basis of present data, in view of the wide ranges of,
~inear velocity variation and the absence of a pressure drop
or energy consumption data base for these data. In addition
to the need for pressure drop.information, the questions of
the relation of the penetration/agglomeration index dependency
to mean particle size and particle size distribution remain
to be answered. .
F-62
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TABLE 4-11
GROSS PENETRATION FOR THE "WETTED SCREEN" SCRUBBER
FOR AN INLET DUST OF MEAl\! DIAMETER 1.5 MICRONS
AND STANDARD DEVIATION 1.8
DATA OF CALVERT AND LEGATSKI (1970)
Concentration Gross
Run Sample {C Penetration
No. (mg./SCM) (gr. /SCF) %
~
I 60 And.. (2) 10..5 0.00459 0.00344 70
m
w 297 11.9 0..00520 0.00390 53
293 13.4 0.00586 0.00439 49
61 And.. 16.1 0.00704 0.00528 48
304 13.9 0.00608 0.00456 49
302 13.9 0.00608 0..00456 50
62 And. 11.0 0.00481 0.00360 54
308 12.5 0..00546 0.00409 62
310 12.9 0.00564 0.00423 51
Air Rate = 2400 CFM
Water Rate = 1.8 GPM
(1) (gr. /SCF)(gal.. /MCF)
(2) Andersen Sampler
-------
100
70
60 .
~
z
o
-
~
0::
I-
W
Z
W
0..
C/)
C/)
o
0::
~
30
20
10
0.00 I
FIGURE 4-13
WETTED SCREEN SCRUBBER
GROSS PENETRATION AS FUNCTION
OF
AGGLOMERATION INDEX. (C d) (WIG)
DATA OF CALVERT 8 LEGATSKI (1970)
,
,0
'0
,
,.
000'
03, o.
,
,
. ,
0.002
0.003 0.004 0.006 0.008 O.O~
. 0.005 0.007 0.00s.
XC' AGGLOMERATION INDEX- (Cd)(W/G)
F-64
-------
E.
PARTICLE SIZE EFFECTS
Meaningful 1itera~ure data on wet scrubber per-
formance as a function of particle size and/or size distri-
bution are very limited. While Lapp1e and Kamack (1955)
tested a group of aerosols of varying mean diameters and
size distributions, as did Krista1, Dennis and Silverman
(1955), comparison of scrubber efficiencies for different
dusts is complicated by the probable secondary effects of
chemical and physical variations. This problem was encountered
by Lapp1e and Kamack, who observed that their ilmenite test
dust showed anomalous behavior when compared to talc dusts,
and found it necessary to separate scrubber performance data
according to the test dust used. What is therefore required
a~ data on the effect of particle size and distribution
variation on scrubber performance for a single dust. The
investigation of Calvert and Legatski on coal dust comes
closest to meeting this requirement, but unfortunately was
necessarily limited to a very narrow range of respirable
dust size. Nevertheless, as is strikingly indicated in
Figure 4-12, an apparently small change in dust size char-
acteristics produced a major shift in the position of the
penetration/Xc line for the Mu1tic10ne. Again, this ;s a
qualitative interpretation, inasmuch as no quantitative data
were provided. Within the restricted limits of the avail-
able data, only a limited projection of the probable size
effects on penetration performance of wet s~rubbers is pre-
sently possible, and this would appear to be an attractive
area for productive research.
1.
Particle Size Distribution
An appreciable part of the problem of relating
particulate scrubber performance to particle size lies in
the proper determination and use of an applicable mean or
representative particle size for a distribution of sizes
in a given dust. There are three mean particle diameter
terms in common use, and these are defined as:
(1)
Number or count median diameter is the
diameter for which half of the total
number of particles have a larger (or
smaller) diameter.
(2)
Mass median diameter is the diameter
for which half the total volume of the
particles is contained in particles
larger (smaller) than this diameter.
F-65
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Specific surface-average or Sauter diameter
is usually applied to liquid drop swarms
and is the diameter of a single drop with
the same ratio of surface to volume as the
total sum of drops.
The relationships between mean particle sizes and
size distributions is covered in standard texts such as Drinker
and Hatch (1954) and will not be detailed here. However, be-
cause scrubber efficiency is generally expressed as weight per
cent, the mass median diameter and mass distribution are the
primary terms of interest. For dusts which follow a logarithmic
normal size distribution, it is possible to convert the count
median diameter, Dc' to the mass median diameter, Dg, by the
relation:
( 3 )
In(Dg) = In(Dc) + 3[ln(ag)]2
(4-3)
where
ag = geometric mean deviation
The obvious sensitivity of Dg to the geometric 'mean deviation
for Dc data in Equation (4-3J imposes stringent requirements
on the analytical distribution data which are difficult to ob-
tain experimentally.
Davis and Truitt (1971) have presented a simple and
significant statement of a normalized mass concentration function
in terms of particle diameter which has been found to fit the
experimental data for many aerosol emissions. This expression
is:
(l/M)dM/dD = [l/ln(Dmax/Dmin)]
(4-4)
where
M = mass concentration
D = particle diameter
max, min = maximum and minimum, respectively
It is believed that this equation has potential utility in
the interpretation of wet scrubbing performance data, and
this will be explored below in connection with the data of
Kristal, Dennis and Silverman (1957).
F-66
-------
2.
Data of Krista1, Dennis and Silverman (1957)
These investigators tested the scrubbing performance
of a Solivore scrubber with a number of aerosols: covering a
wide range of mass median particle sizes and inlet dust load-
ings. Penetration test data were obtained for several aerosols
at varying water rate runs, and for a separate series of vary-
ing inlet dust loading runs; these data are presented in
Tables 4-12A and 4-12B. Before analyzing these data in terms
of the agglomeration index, it is necessary to review the
methods of preparing and introducing the aerosols, inasmuch
as these definitely affect the results.
The coarse H2S04 mist was prepared by aspiration,
followed by removal of the larger particles by impingement on
a baffle in an e1utriation chamber. Similarly, the CuS04 was
produced by pneumatic nozzle atomization, followed by elutriation,
and finally dehydration in a heated pipe. The fly ash, CaC03'
and other solids were fed to the scrubber by means of the
"Harvard generator" for loadings greater than 0.2 gr/CF, and
the National 8ureau of Standards gear-tooth/aspirator feeder
was used for loadings in the range of 0.02 gr/CF. As noted
earlier, the use of the latter two types of feeder involve
the possibility of introducing pre-agglomerates or forming
dry agglomerates in the venturi aspirator. The secondary
e1utriation and/or impingement treatment of the CuS04 and
H2S04 aerosols ensures the absence of agglomerations in the
scrubber feed, and Figure 4-14 treats these two aerosols as
comparable.
Figure 4-14 indicates that the limited penetration
run data for the two aerosols fit the -450 agglomeration index
line relationship previously found for both Ingels' data on
asphalt mix plant dust, and Calvert and legatski's coal dust
penetrationsr The agreement between the two sets of aerosols
tests in Figure 4-14 is surprisingly good, and a single line
could obviously be drawn for both sets of points rather than
the individual lines actually shown.
While the CuS04 and H2S04 aerosols were prepared
by virtually identical methods, the fly ash and CaC03 mat-
erials in the Table 4-12A and 4-12B runs were fed to the
scrubber by differing means. The data for the latter dusts
are plotted in Figure 4-15, and the anomalous behavior, com-
pared to Figure 4-14, is apparent. The line through the two
CaC03 points has a slope of -0.4, indicating a possible pre-
agglomeration effect with either one or both of the two dif-
ferent f~eders used for the two points.
F-67
-------
TABLE 4-12A
MUL TI - VENTURI (SOLIVORE) SCRUBBER
PENETRATION TEST DATA FOR TIiREE AEROSOLS
AT VARYING WATER RATES
DATA OF KRISTAL. DENNIS AND SILVERMAN (1957)
Aerosol WIG Inlet Dust (Cd>(w /G) Weight Passage
gal/1000 Loadin~ Collection %
ft.3 of air gr ./ft. Efficiency %
Fly Ash 20 002(1) 400 99.2 0.8
(14.3 30 002 6.0 99.4 0.6
micron) 40 0.2 8.0 99.6 0.4
CuS04 20 0.6/1000(2) 00012 78.9 21.1
. (0.74 30 0.6/1000 0.018 85.0 15.0
micron) 40 0.6/1000 0.024 90.2 9.8
H2S04 20 2.5/1000(3) 00050 9208 . 7.2
(13.8 . 30 2.5/1000 0.075 95.5. 4.5
micron) I 40 2.5/1000 0.100 96.2 3.8
TABLE 4-l2B
PENETRATION TEST DATA FOR VARYING INLET DUST LOADING
Aerosol 'W/G Inlet Dust (Cd>(W /G) Weight Passage
gal/1000 Loading Collection %
ft.3 of air gr ./ft. 3 Efficiency %
Fly Ash 30 0.02(4) 0.6 99.0 1.0
(14.3 1.60(1)
micron) 30 48.0 ' 99.4 0.6
CaC03 30 0.25(4) 7.5 88.2 11.8
(2.6 i
I
micron) 30 1.50(1) 45.0 93.4 6.6
(1)
(2)
Harvard generator feeder
Aspiration/Elutriation preparation
(3)
(4)
Aspiration/Impingement preparation
NBS feeder
. F-68
-------
.~
z
o
t-
«
cr
t-
W
Z
lIJ
0-
Ct)
Ct)
o
a:
(!)
100
80
60
20~
10
8
6
1.0
QOI
FIGURE 4-14
SOLIVORE' SCRUBBER
PENETRATION AS FUNCTION
OF
AGGLOMERATION INDEX, (Cd) (WIG)
DATA OF KRISTAL,ET AL (1957)
~
. ~'o ./"SLOPE --I
COUNT "V
MEDIAN
DIAM.7 M
O-CUS04 0.48
a-H2so4 4.00
0.02
0.04 0.06 0.10
0.08
0.2
0.4 0.6 0.8 1.0
. XCI AGGLOMERATION IN DEX. (Cd) (WI,?,)
F-69
-------
~
z
o
~
-------
The fly ash data of Figure 4-15 exhibit several ob-
vious anomalies, in addition to any that may have been caused
by feeder variation. The central three points, from Table
4-12A for the varying water rate series of runs, appear to
fall on a -450 line, parallel to the other two run series in
this sequence plotted in Figure 4-14, despite the difference
in feeder type. However, when the varying inlet dust loading
data of Table 4-128 are added, complete data disagreement re-
sults. The data point on an Xc value of 0.6 was obtained with
the NBS dust feeder at a dust loading of 0.02 grlCF, while
the data point at Xc = 48, using the Harvard feeder, corres-
ponds to a dust loading 80 times greater, or 1.60 gr/CF. In
order for these points to be aligned with the -450 line drawn
through the three central points, the penetration value at
Xc = 0.6 would have to be 5.4% and that for Xc = 48, 0.07%.
While pre-agglomeration prior to scrubber entry may effectively
decrease the penetration from the anticipated 5.4% to the.
actual 1.0%, it cannot explain the increase in penetration
from the expected 0.07% to the actual 0.6% at the other end
of the line. There appears to be a probable explanation for
this effect, however, in the analytical data provided by the
authors and given in Table 4-13.
The use of a settling bottle in the effluent sampl-
ing line enabled the collection of wetted dust agglomerates
which escaped collection in the scrubber, thus yielding the
two sets of penetration values listed in Table 4-13. It was
not stated whether the data of Tables 4-12A and 4-128 were
given on a dry or wet basis~. but the fly ash run of Table 4-13
corresponds to an Xc value of 4.5, and yields penetrations .of
0.6% dry, and 1.6% wet. The dry value falls close to the
-450 line of Figure 4-15, while the 1.6% penetration is higher
than any of the other fly ash data points. The functional ef-
ficiency of a settling bott1e.in removing the wetted dust ag-
glomerates is unknown, and it is probable that a fraction of
the wetted-agglomerate reported to the dry sample, which could,
at the sensitive low end of the penetration scale, easily ac-
count for the displacement of a penetration value from 0.07%
(projected) to 0.60% (indicated). As discussed previou~lyin
terms of the data of Calvert and legatski, the experimental
difficulties in accurately determining penetration levels
less than 1.0% undoubtedly contribute to the apparent dis-
crepancy.
The appearance of wetted-dust agglomerates in the
scrubber effluent samples for the Solivore runs parallels
the behavior observed by. Ingels for the multi-stage centrifugal
F-71
-------
TABLE 4-13
EFFECT OF SAMPLING METHOD ON ESTIMATION OF EFFLUENT LOADING
SINGLE -STAGE SOLIVORE SCRUBBER
DATA OF KRIST ALp DENNIS AND SIL YERMAN (1957)
Outlet Loadings Efficiency
Inlet
Loading Dry Wetted Total Dry(a) Wet(b)
Aerosol gr./CF gr 0 /CF gr 0 /CF gr 0/ CF % %
Fly Ash 0.15 0.0009 0.0015 0.0024 99.4 98.4
Talc 1.5 0.056 0.027 0.083 96.2 94.4
H2S04 0.0025
(coarse) mist(c) 0.00011 0.00027 0.00038 95.5 84.7
(a)
"Dry" refers to dust passing settling bottle in sampling line.
"Wet" refers to dust entrained in water droplets retained in settling bottle.
(b)
(c)
No droplet eliminator in collector.
F-72
-------
scrubber, discussed earlier. It is obvious that in both cases,
the scrubbers are not doing an adequate job of removing the
dust-water agglomerates formed, even though the agglomerate-
formation contacting efficiency is high. The very fact that
the contacting efficiency is high may impose too great a de-
mand on the removal function of the unit. Inasmuch as the
agglomeration index correlation approach assumes the removal
from the gas phase of all agglomerates of a critical minimum
size, data for equipment utilizing a separate positive mech-
anism for agglomerate removal, such as a venturi-cyclOne com-
bination, should correlate better than data for a unit provid-
ing inefficient removal. The latter defect in scrubber de-
sign will be evidenced by excessively high penetration levels
at high X values, and the agglomeration index thus becomes a
measure of the adequacy of the specific scrubber design function
of agglomerate removal.
3.
Size-Dependent Penetration Behavior
Because of the limitations of data availability,
it is difficult to go beyond a speculative level of analysis
of penetration dependence on particle size. Nevertheless
there are a number of suggestive patterns of behavior in
the penetration graphs presented to warrant preliminary dis-
cussion. In Figure 4-14 the agreement in penetration behavior
on an agglomeration index plot between aerosols of widely
different particle sizes raises some fundamental questions.
The respective count median diameters of the CuS04 and H2S04
microspheres were 0.48 and 4.0 microns, so that, on a pop-
ulation basis, there were roughly 10 times as many CuS04
particles per unit volume as there were H2S04 mist droplets,
at equal Cd levels.
If agglomeration with water drops or self-agglom~
eration of the dust to critical removal size is particle pop-
u la t i on - de pen den t , as a 11 of the ex per i menta 1 data i n d i cat e ,
then the fractional penetration of the CuS04 aerosol should
be substantially lower than that of the H?S04 mist at the
same run conditions. However, if the limlted data of Figure
4-14 are truly representative, the performance 01 the scrubber
for CuS04 is substantially identtcal to that for H2S04.
There are several explanations for this behavior. On the
basis of conventional impaction theory, the target efficiency
of the water drops for the smaller-size fractions of the
CuS04 aerosol is probably negligible, and certainly less
than for the H2S04' which would tend to offset the effect of
the higher particle number concentration. Thus, the agree-
ment of the penetration behavior of the two aerosols may be
regarded as fortuitous, resulting from a coincidental balance
of opposing capture forces.
F-73
-------
An alternate explanation, and one with fundamental
significance, is afforded by the recent work of Davis and
Truitt (1971). These investigators calculated atmospheric
aerosol concentration as a function of size distribution for
a steady-state aerosol injection rate in balance with a tur-
bulent-mixing IIcoagulationll and gravitational fallout rate.
It will be noted that this is essentially the same general
mechanistic scheme proposed as the basis for scrubber oper-
ation in this report. In theii work, Davis and Truitt as-
sumed an input particle mass distribution in accordance with
Equation (4-4), with limits of Dma¥=lOO microns, and Dmin as
the independent variable, ranging Trom 0.01 to 10 microns~
For the assumed values of D i ' the steady-state airborne
concentrations were calcula~eg for removal by coagulation
and settling, with the results given by Figure 4-16. The
highest steady-state concentration was found to result from
minimum particle diameters of 0.2 to 2 microns.
If the analogy can b~ fully drawn between the Davis
and Truitt atmospheric fallout model and the action of a wet
scrubber on an accelerated time-scale, then the residual dust
penetration (equivalent to the steady-state airborne concen-
tration) would be a function primarily of the minimum particle
size, assuming the mass-size distribution of Equation (4-4)
and a"fixed Dm x. Qualitatively, the shape of Figure 4-16
was explained gy Davis and Truitt in a manner that would also
apply to wet scrubber action. If the minimum particle size
is relatively large, say, greater than 2 microns, then all
of the particles are readily removed, and penetration would
be quite low. This situation corresponds to the scrubber
case where agglomeration action is not necessary, and the
agglomerate-removal mechanism in and of itself is sufficient
for efficient operation. Alternately, if the minimum particle
size is quite small, say, less than 0.2 micron, then the
number concentration is very high and the agglomeration (co-
agulation) rate is high enough to yield rapid formation of
particles greater in size than the critical removal size. In
other words, for small particles, agglomeration governs, while
for large particles, the removal mechanism controls penetration.
The intermediate particle size range, which for Figure 4-16
would be 0.2 to 2 micron, constitutes the most stable or dif-
ficultly-removable size fraction for the atmospheric holdup
problem.
With reference to the atmospheric particle stability
problem, the comment of Davis and Truitt regarding th~ stable
size range is of pertinence her~: 11$0 in summary, the particle
F-74
-------
10.1
-
..,
E
.......
CP
-
z
o
~
-------
size range (perhaps roughly 0.2-2 micron diameter) constitutes
the worst of particulate air pollution. These are the particles
that get deposited in the lungs; these are the particles that
cause reduced visibility;'these are the particles that settle
slowly and have a long airborne life. Indeed, it appears that
particulate air pollution could well be defined as the 0.2-2
micron particles.1I
Some, rather remarkable prior experimental confirm-
ation of the Davis and Truitt theoretical treatment is given
by the study of Langstroth and Gillespie (1947) of coagulation
and surface losses of smokes in still and turbulent air. Re-
sults of the particle size distribution studies 6n ammonium
chloride smoke showed that in still air, the coagulation and
fallout sequence gave a stable residual particle size of 0.8
to 1.2 micron, from the smoke having an initial size frequency
peak in the, 0.01 to 0.2 micron range. Turbulent or stirred
air gave a stable size frequency peak corresponding to the 0.05
to 1.0 micron range. However, in the latter case, the size-
frequency peak shifted with time from a value of about 0.03
microns at 100 minutes'to 0.07 microns at 280 minutes indicat-
ing the accelerative effect of turbulence on coagulation. 'The
similarity between the experimental curves of Langstroth and
Gillespie and the theoretical curve of Davis and Truitt in
Figure 4-16 is quite striking, and the possible significance
of these data to wet scrubber work is worth exploring.
The question here arises as to what extent turbulent
coagulation-fallout theory actually applies to wet scrubber
performance. The preliminary indications given by the agglom-
eration-index correlation work of ,this design review are that
wet scrubbers may very well obey the same rules apolying to
the more extended time-scale atmospheric dust senaration nhe-
nomena, and coagulation-removal theory may afford a better ex-
planation for scrubber data than imDaction theory. Certainly,
the assumption that wet scrubbers are merely accelerated time-
scale turbulent mixing devices yielding the same results as
the atmospheric process serves to explain some well-known and
seemingly anomalous scrubber characteristics. For examDle, it
has been shown by Kristal, Dennis and Silverman (1957) among
others, ~hat for several scrubbers or scrubbing stages in series,
penetratlon wlll follow the log-penetration law:
ET = 1 - (1 - Eo)n
= 1 - (Fo)n
(4-5)
(4-6)
F-76
-------
where
ET = overall scrubber efficiency, fractional
Eo = stage efficiency, constant
Fo = penetration of a single stage
n
= number of stages
Equation (4-5) properly applies only to the collection of uni-
form aerosols, and as Kri~ta1 ,et a1 notes, lIin order for the
second and third stage of the experimental collector to be as
efficient as the first stage, the effluent from any preceding
stage must have undergone sufficient conditioning (through ag-
glomeration, particle and water contact, and possibly condens~
ation) to approach the size distribution of the original aerosol.
In the absence of particle conditioning, multi-stage operation
would ordinarily be impractica111. The fact that the experimental
data for multi-stage scrubber collection follows the log-pen-
etration statement of Equation (4-5) has been difficult to
reconcile with the concom1ttant necessity of accepting con-
stancy of particle size distribution, as per first stage,
throughout the equipment, despite the removal of 98% or more
by weight,of the aerosol. It is obvious that Figure 4-16, and
the supporting turbulent-agglomeration atmospheric fallout
theory, supplies the answer. Given the difference in turbulent
mixing intensity, residence time and removal mechanism of wet
scrubbers, the stable size fraction is probably higher than
that shown in Figure 4-16, and literature data suggest that it
may lie in the 2 to 20 micron range. Nevertheless, turbu~ent
agglomeration and surface or inertial removal in the first
stage should produce a IIstab1ell size distribution within that
stage which should persist in stability throughout all stages
as long as the mechanisms of agglomerative contacting,and ag-
glomerate removal remain the same. While the constancy of
particle size distribution through repetitive scrubbing stages
may be difficult to accept, this behavior has been unavoidably
inherent in the experimentally-verified log-penetration law
behavior of multi-stage units.
A further insight is provided by the observation of
Langstroth and Gillespie (1947) that the rate of coagulation
depends on the square of the number of particles, while the
rate of sedimentation (or~inertial removal) depends on the
first power of the particle concentration. This suggests that
it would be possible to test a scrubber for primary function
by making a series of varying concentration runs at several
fixed particle size distribution, and determining the relation
F-77
-------
between penetration and particle c~ncentration. This is es-
sentially what was done in some of the dry runs reported by
Calvert and Legatski, and the dependence of penetration on
the -1 power of Cd for such units as the Mu1tic10ne and packed
tower would indicate that these are non-agglomerating, pure
inertial removal devices. It should also be pointed out that
the agglomeration index treatment for wet scrubbers is pre-
dicated on the product of the water drop and dust particle
concentration terms (replacing the square of dry dust concen-
tration) and assumes that dust-dust collisions do not contri-
bute significantly to agglomeration. All of the assumptions
and findings of this report are subject to direct testing in
the context postulated, and such a program is obviously in-
dicated. In the meantime, it is informative to review ad-
ditional data which does not fit the r~lationships pr~vious1y
established for penetration behavior in wet scrubbers.
F.
ANOMALOUS BEHAVIOR
In addition to the above-noted Krista1 . et a1 (1957)
data on CaC03 and fly ash, and the fly ash data of Johnson
(1955) reviewed earlier, departure from the expected inverse
relationship between penetration and dust loadings, or the
agglomeration index, is observed in the data of Lapp1e and
Kamack (1955) and Lancaster and Strauss (1971). .
1. Data of Lapple and Kamack (1955)
The experimental studies reported by Lapp'le and
Kamack covered performance tests on the effect of operating
variables on the efficiency of several laboratory and semi-
works scale scrubbers. The important variables were indi-
cated to be inlet dust loading, water/gas ratio, gas velocity
and/or pressure drop. While the authors' final conclusion was
that contacting efficiency, at constant inlet dust loading,
is solely a function of the total gas power'input (degree of
turbulent mixing) the effect of varying (WIG) ratio on pen-
etration was shown only by independent points for this vari-
able, and cross-correlation was not attempted. Replotting
the lapple and Kamack d~ta on a (Cd)(W/G) basis, rather than
dust loading alone, eliminates the effect of varying (W/G)
and results in a single series of constant-gas velocity lines
for each scrubber. However, the data are limited and not very
consistent, with the ~lopes of the penetration vs. agglomeration
index lines varying from -0.1 to -1.0.
F-78
-------
The erratic behavior of their penetration-dust load-
ing plots was noted by Lapple and Kamack, and explained on the
basis of pre-agglomeration of the dust, and directly attributed
to "a failure of the dispersing venturi to disperse completely
the bulk dust". This effect was particularly apparent for the
ilmenite and titanium oxide tests dusts, and much less so for
the two talc dusts used. Although only two points were pre-
sented for the specific low dust concentration region, the
authors indicated that the penetration approached a "constant
value" for dust loadings less than 2 gr/CF, but decreased rapidly
at loadings above 10 gr/CF. Actually, this pattern is occasion-
ally quite evident in an agglomeration-index treatment of their
data, such as Figure 4-17. The explanation advanced by Lapple
and Kamack for this effect of dust loading was that lithe dust
present at high concentration is effectively coarser because of
agglomeration than the same dust at low concentration", even
though the particles were statistically 20 to 80 diameters apart
at the higher concentrations.
This argument that the degree of pre-agglomeration is
concentration-dependent is refuted by the data of Ingels and of
Calvert and Legatski reviewed earlier. The latter test systems
contained a preliminary inertial separator between the dust
source and the test scrubber. This type of arrangement serves
to remove the agglomerates formed by the dispersing unit, or
which remain undispersed from the bulk dust, and yet tests car-
ried out under these conditions clearly show a constant pene-
tration/dust loading relationship which is invariant with dust
concentration range. Thus, the Lapple-Kamack hypothesis of
pre-agglomeration in the feed dust only at the higher concen-
trations does not appear to be supportable. In fact, the more
likely explanation is the reversal of that advanced: pre-ag-
glomeration causes the flattening of the curve of Figure 4-17
in th~' low Cd range, rather than the high Cd range. As stated
earlier, if pre-agglomerated material is fed to the scrubber,
then the apparent efficiency will be higher than that to be
expected if the same dust loading consisted of completely dis-
persed particles. The agglomeration function normally required
of the scrubber has already been partially performed, and the
test is really of the scrubber agglomerate-removal function.
Thus, penetration will be lower than normal, with resulting de-
pression of the usual curve. The fact that the curve increases
in negative slope with increasing dust loading, approaching
the normal -1 value, indicates either that the turbulent-mixing
mechanism of the scrubber is causing additional re-dispersion
of the dry-dust agglomerates, or that the fraction of ore-
agglomerated material diminishes at the higher concentrations.
F-79
-------
FIGURE 4-17
PERFORMANCE OF I-INCH S-SEND CONTACTOR
AND a-INCH CYCLONE IN'SERIES
, ' '
, ":PENETRATION' AS FUNCTION OF '
, . AGGLOMERATION INDEX
DATA OF LAPPLE a KAMACK (1955)
10
8
6
4
~ . ----~~
z 2
o --0 ~
-
t- . 0 0,
oCt
a:
t-
UJ 1.0
z'
UJ 0.8
a.
00' 0.6
en
0
a: SLOPE.-I
'0' 0.4
02
0- TALC IIAI} ,
II II VELOCITY D 98 FPS
+-TALC B ,
", I . .
0.1
0.1
0.2
0.4 0.6 0.8' 1.0 "
2
4
6 . 8 10 '
. XC' AGGLOMERATION INDEX .'(Cd)(W/G)
F-80
-------
2.
Data of Lancaster and Strauss (197l)
The influence of the pre-agglomeration effect on
scrubber performance is still occasionally unrecognized.
The recent study of Lancaster and Strauss (197l) on the in-
fluence of steam injection in a wet cyclone scrubber pre-
sents a series of penetration-loading curves which are re-
markably similar to Figure 4-17, even with respect to the
loading range of 2 to 5 grlCF at which the curves IIbendll
toward a slope of -1. These curves are for runs both with
and without steam injection, so that the effect is cnar-
acteristic of the f~ed/scrubbe~ system. Inspection 6f the
test setup for this study shows that the IIHarvard generatorll
and II d i s per sin g II vent u ria r ran gem en t was use d, and i tis
apparent that these data suffer from the acknowledged de-
ficiency of the Lapple and Kamack data, that of feed pre-
agglomeration. It is unfortunate that this fundamental
anomaly continues to be i"troduced into scrubber test work,
inasmuch as it is quite obvious that without this invalidat-
ing factor, performance and design correlation efforts would
be much further advanced than they are at present.
G.
PENETRATION AS A FUNCTION OF SCRUBBER PRESSURE LOSS
The third primary scrubber variable indicated by
the literature on scrubber performance is the gas velocity
or pressure drop. Lapple and Kamack made extensive measure-
ments of the effect of this variable on penetration (un-
fortunately at various levels of feed pre-agglomeration)
and concluded that the '~ontrolling factor in scrubber per-
formance must be one associated with pressure drop. Such
a factor is turbulencell. Unlike the apparently simple
relation of gross particulate penetration to the other two
primary scrubber variables of dust and water drop concen-
trations, the effect of pressure loss on penetration is
complex. On the basis of the turbulent agglomeration mech-
anism which appears to govern the observed wet scrubber be-
havior, it would be expected that there would be two dif-
ferent zones of penetration dependency on gas flow, corres-
ponding to (a) laminar flow and (b) fully-developed tur-
bulent flow. This is approxf~at~ly.the behavior observed
in full-range flow experiments on wet scrubbers.
The most complete data on pressure drop are those
of Lapple and Kamack, and a typical curve for penetration as
a function of scrubber ~P is given in Figure 4-18 for the
Venturi orifice or cyclone scrubber. In general, the typical
curve consisted of three regions: a low ~P region of'less
F-8l
-------
F"IGURE 4-18 .
. . . - .
. .
.' "
TYPICAL BEHAVIOR OF PENETRATION
. . I .
. .
AS FUNCTION OF PRESSURE DROP
DATA OF LAPPLE a KAMACK (1955)
100
..
Z 3
o
-
t-
<{
a:=
I-
UJ
Z
UJ
Q.
0.7
0.1
I
2
3
5 7 10 20 30 50
PRESSURE DROP, INCHES H20
F-82
-------
than 7" W.C., where the log-log penetration/flP plot had a
slope of about -0.5, a high flP region above 15" W.C., where
the slope was approximately -1.8,and an intermediate zone
from 7 to 15" flP connecting the low and high flP regions.
This zone may either be drawn as a curved segment, or a
straight line of -1 slope. This curve was obtained for
all dispersed-water contactors tested except for the cases
of the sieve tray, dry runs, or for seriously pre-agglom-
erated feed dusts. The model of water introduction and its
initial degree'of dispersion had only a small effect on the
relative position of the dust 10ss/flP curve, and, aside from
the exceptions noted, the type of contactor did not cause a
major shift in curve location.
The standard sieve tray runs produced a straight-
line penetration/flP relationship, with a slope of -0.75,
confirming the finding of the Calvert and Legatski data
that this unit does not utilize the normal dispersed-water
contacting mechanism. The pre-agglomeration effects in the
ilmenite and titanium dioxide runs on the standard penetration
plots were drastic, with increases in penetration occuring
with increasing velocity (flP). By manipulation of the water
rate and equipment type~ it was possible to obtain a limited
amount of varying flP data at constant inlet velocity con-
ditions for these problem dusts, and these data gave normal
plots of the type shown in Figure 4-18. However, these con-
stant-velocity curves are displaced to the right with increas-
ing gas velocity, and it is uncertain whether this effect is
the result of a more complete degree of dispersion of the
entering feed dust or a separate secondary effect of gas
velocity in water atomization. It was visually observed
that, for the Venturi and pipeline scrubbers, a marked im-
provement in the degree of water atomization took place as
the pressure drop approached and exceeded 15" W.C., but Lapple
and Kamack considered this to be a minor secondary effect, and
indicated that the level of gas turbulence controlled. Again,
it should be noted that, without the pre-agglomeration com-
p 1 i cat ion ,. i two u 1 d h a v e bee n far 11 y s imp 1 e to s e par ate and
identify the effect of water atomization as a function of gas
velocity by means of the constant velocity-varying flP runs.
The linearity of fractional dust penetration with
flP above 15" W.C. obviously enables a complete and simple
statement of the relation of penetration to the three pri-
mary variables for this range:
F = K(flP)-1.8
(Cd)(W/G)
(4-7)
F-83
-------
where K is a performance constant, characteristic of the
scrubber. and the sum of the secondary design variable contri-
butions. However, even if data are obtained below the linear
high ~p range of Figure 4-18, knowledge of the full-range
behavior facilities interpretation and understanding of the
penetration results. A case in point' is the work of Byrd and
Dewey (1957) on application of a Venturi scrubber to the re-
moval of odor attendant on the emission of submicron product
fines and liquid aerosols. The Byrd and Dewey data were pre-
sented as smoothed-curve plots of percent odor removal and
Venturi pressure drop as functions of (WIG), for constant
parametric values of gas velocity through the Venturi throat.
By cross-plotting these data, it is possible to obtain graphs
of percent odor penetration vs: scrubber ~p at constant gas
velocity, and the Byrd and Dewey data are presented in this
form in Figure 4-19. Although subjectively measured by panel
testing, the "quantitative" odor penetration data appear to
follow the same type of ~p curve established by Lapp1e and
Kamack for the dry particulates. While the data of Figure
4-19 are indirect, and are interpolated from the original
smoothed curves, they do provide some insight into the
probabl~ separate effects of gas turbulence and water dis-
persion.
As shown in Figure 4-19, the Byrd and Dewey data
fell into two groups, depending on the range of the gas vel-
ocity. The upper curve of Figure 4-19 covers gas velocities
to 218 ft/sec, and this group of points form a line that is
consistent with the behavior of the Lapple and Kamack data
of Figure 4-18, although the range of ~p is limited to 17"
w.c. However, the group of data points corresponding to
Venturi throat velocities of 242 to 338 ft/sec appear to
fall on a separate curve of considerably lower (about one-
third less) penetration levels. These data definitely in-
dicate a separate effect of gas velocity on water atomization,
in addition to the turbulent-mixing gas velocity effect
directly measured by ~P. Further, the apparently "critical"
gas velocity classifying the data into two groups corresponds
to the critical range of gas velocities above which water
drops will shatter, approximately 200 ft/sec. Above this
minimum, the drop size is controlled by gas velocity-induced
atomization, and the drop size -produced in a Venturi is
9iven by the empirical equation of Nukiyama and Tanasawa
(1938), which for water systems can be stated (Johnstone,
1949) as:
Do = l6,000/Vg + 1.4(W/G)1.5
(4-8)
F-84
-------
100
80
60
cf!.
Z 40
o
t-
-------
where
Do = mean Sauter drop size, microns
Vg = relative air-water drop velocity, ftlsec
As indicated by Equation (4-8), drop size, and there-
fore drop population, becomes gas velocity-dependent in the
plus-200 ftlsec range, and secondary dependencies on (WIG) and
V are introduced. The behavior of the Byrd and Dewey data is .
t~us in accord with estal)lished' dispersion effects for Venturi.
scrubbers, and because of the absence of any pre-agglomeration
effects for this data (a pre-cyclone scrubber was used) it is
probable that the Figure 4-19 separation by velocity range is a
more accurate description of Venturi penetration performance
than Figure 4-18. While Figure 4-19 does show self-clasSiftcation
of the penetration data into pre-critical and above-critical gas
velocity ranges, insufficient data are available in the ,above-
critical zone to permit differentiation of the velocity lines
in accordance with Equation (4-8). The additional (WIG) de-
pendency introduced by the drop-shattering phenomenon is a rel-
atively small correction to DQ in the 200.400 ftlsec velocity
zone, and would not be signif1cant at values of (WIG) less than
10 gal/MCF.
The data of Figures 4-18 and 4-19 conceal some in-
herent contradictions which are well worth analyzing in con-
junction with the use of EquatiDns (4-7) and (4-8). The
slopes of the sets of velocity lines in Figure 4-19 are sub-
stantially -1 in the range above 7" W.C.. This is essentially
the behavior of the Lapple and Kamack data of Figure 4-18 for
the 7" to 15" ~p range, and this indicates that penetration
is inversely propartional to the ~p of the scrubber. If, as
Lapple and Kamack concluded, and as the data reviewed in this
report have demonstrated, penetration is controlled by the
degree of gas turbulence, then penetration should be inversely
proportional to ~p throughout the range of fully-developed
turbulence including the +20" region. However, Figures 4-18
and 4-19 show that the -1 relation exists only up to about 15"
to 20" W.C. ~P, and then fairly abruptly breaks off to the
-1.8 slope. Ignoring for the moment the additional ~p ef-
fect that can be attributed to incremental liquid holdup and
flow resistance caused by increases in (WIG), the most likely
explanation of the increased slope of the penetration/~P
plot in this range is that the efficiency increases because
of the incremental effect of water atomization at this in-
creased turbulence level. Equation (4-8) states that the
mean drop diameter is inversely proportional to the gas velocity
(square root of ~P) at constant (WIG) levels, and if water
F-86
-------
droppopulation number is inversely proportional to the cube
of the drop diameter, then as a first approximation:
N = (k/Do)3 = (k'Vg)3
= k"(f1p)leS
( 4- 9 )
(4-10)
where
N = water drop population concentration
Equation (4-10) indicates that penetration should
be inversely proportional to (f1P)2.S in t~e water atomization
region, rather than the 1.8 power shown in Figure 4-18. How-
ever, in Equation (4-9) the correct drop diameter required
is the count mean diameter rather than the Sauter mean drop
diameter, so that more properly, the slope of the f1P line
in the +1511 W.C. region in Figure 4-18 should only be ex-
pected to be in the -1 to -2.5 range. Conversely, the in-
crement of -0.8 in the slope actually obtained as a result
of water atomization would call for re-writing Equation
(4-10) as: -
N = kll(f1P)O.8
'(4-1l)
The limited range of f1P in the Byrd and Dewey data
do not permit examination of the validity of Equation (4-11)
for the water atomization region, but additional information
on the high-range penetration/f1P behavior in a Venturi scrubber
is provided by the work of Brink and Contant (1958) on scrubbing
of phosphoric acid mist. Mist loadings were held constant
at 1400 mg/CF of P20S' and factorial design of the experiments
enabled studies of t~e individual variables of spray liquid
rate, throat gas 'velocity, liquid injection velocity, and
total number of spray jets. The water jet spray velocity
was found to significantly influence the performance of the
Venturi, with velocities greater than 25 ft/sec yielding
higher penetrations. In replotting the Brink and Contant data,
it was found that using the higher spray velocity data of 30
and 40 ft/sec produced scatter and generally higher penetration
values than were characteristic of the low-velocity jet runs,
and accordingly, log-log penetration/6P plots were made only
for the latter data. A typical plot of the Brink and Contant
data is presented in Figure 4-20 for percent penetration vs.
overall scrubber pressure drop for an injection rate of l!IT
GPM. A straight line with a slope of -2.0 is obtained from
the data, and it should be noted that the f1P range is 20-4011
W.C. This latter range corresponds to the high-f1P segment
F-87
-------
~
o
...
~4
-
.-
-------
of the Lapp1e and Kamack curve of Figure 4-18, and the agree-
ment between the -2.0 and -1.8 slope values may be considered
to be fairly good.
Plotting additional Brink and Contant data at an in-
jection rate of 200 GPM gave a line parallel to that of Figure
4-20, but at the expected lower penetrations. It was believed
that the water jet velocity effect caused by transverse in-
jection noted by Brink and Contant was due to a momentum ex-
change which effectively lowered the relative gas velocity.
If such was the case, then at constant water'jet velocity,
the initially higher penetration observed at the superficial
velocity as measured by the ~P should decrease with increasing
~P. This was confirmed by several of the graphs made for the
high-velocity jet runs, which showed curved lines approaching
the Inorma1" low-velocity jet line at the high ~P values, but
data scatter was too great to generalize this effect.
The analysis presented above considers penetration
to be inversely proportional to both the degree of turbulent
mixing and the population concentration of the water drops
(at constant Cd). The agglomeration, or dust-water collision
concept, postulates that the intensity of gas turbulence and
population concentration of water droplets are separate and
independent primary variables. Inasmuch as each of these two
variables is individually dependent on the first power of ~p
in the drop-shattering range, the concept qualitatively de-
scribes the observed ~P behavior of wet-scrubbers. Addition-
ally, below the critical range of gas velocities which cause
drop-shattering, the penetration should be dependent on the
-1 power of ~P, rather than the -2 power of the higher ~P
range, and again, the observed behavior is in accord with these
expectations.
There is reason to suspect that, in terms of ~P
behavior. Venturi scrubbers may show characte~istics not
exhibited by other wet scrubbers. As Wink1ep1eck (1970)
has shown, for the Venturi, a conservation of energy balance
yields an expression for scrubber pressure drop in terms of
gas velocity, gas density, and liquid injection rates;
~P = Vg2[Apg + B(W/G)]
(4-12)
where
A. B = constants
Thus, a given ~P may be obtained by adjusting either the
gas velocity or liquid rate, and there will be an infinite,
F-89
-------
number of combinations of these latter variables providing
the given 6P. Equation (4-12) indicates that it may be more
rational to work from a velocity and liquid/gas ratio base
toward an efficiency correlation for Venturi units, but such
separate data are seldom available.
It is obvious that, while the turbulent-agglomeration
approach is rational and empirically productive, it lacks an
adequate supporting theoretical base. Thus, one of the prime
objectives of any research and development program based ~n
the correlation of this report should be the development of a
consistent theory.
. F-90
-------
v.
CORRELATION IN SEARCH OF A THEORY
The agglomeration-index correlation presented in
Section 4 has been empirically-derived, but it has the
singular advantage of fitting wet scrubber experimental and
field data. Anomalous data that do not follow the cor-
relation are readily accountable in terms of known inherent
deficiencies in either the test methods or the scrubber
itself. The final correlating equation:
K(~p)-n
F = (Cd)(W/G)'
where n = integer of 1 or 2, depending on
~p regime
is usable in its present form both for scrubber performance
correlation and for design. Equation (5-1) allows, for the
first time, valid and absolute performance comparisons be-
tween scrubbers of different types, and establishes the
ground-rules for parameter constancy which must be observed
to validate such comparisons. It further provides a basis
for direct parametric testing and research on wet scrubbers,
and the need for programs of investigation of the effect of
particle mass and number distribution and other variables
in context has already been noted. The one definite and
acute deficiency is the present lack of a theoretical base,
which persists despite several preliminary attempts at
reconci1iati6n of ,inertial impaction theory with the data
treatment of the correlation approach. It has gradually
been realized that not only does impaction theory fail to
support the empirical correlation of this report, it vails
to correctly describe fundamental wet scrubber behavior itself.
(5-1)
All of the correlations derived have consistently
indicated ,that penetration correlates with such variables as
(WIG), AP; and Cd on log-log plots. On the other hand, im-
paction theory calls for a semi-log relationship between
penetration and WIG (or ~P) and obvious~ only one of the two
possibilities can be correct. A search of the literature
showed that a direct experimental data test of the semi-log
relationship between penetration and (W/G)(v)lh~ as a function
of particle size was made by Brink and Contant (1958) and
Johnstone, Feild and Tass1er (1954) who produced a "linear"
function for their data on such coordinates. However, even
a casual inspection of the original plots shows a pronounced
misfit of the experimental points~ which exhibit a distinct
curvature despite the imposed' straight line. The Brink and
Contant data have been replotted on log-log coordinates and
are presented in Figure 5-1.
F-91
-------
1.0
it
z 0.07
o
t= 0.05
q:
a:
~ 0.03
z
L\I
~ 0.02
0.01
0.002
0.001
. I
FIGURE 5-1
PENETRATION AS FUNCTION OF
MbDIFIED INERTIAL PARAMETER
DATA OF BRINK a CONTANT (1958)
\
3
5
50 70 100
2
7 10
w.r;r
20 30
F-92
-------
The fit of the data points to the straight line of
Figure 5-1 is far better than ih'the original semi-log co-
ordinate plot, and is consistent with the behavior of other
(WIG) data examined in this report, except for the higher
slope of Figure 5-1. In the Brink and Contant inertial
impaction treatment, the abscissa values were the product
of the 1iquid/¥~s' ratio and the dimensionless inertial
parameter, (~) I~, and this grouping has been retained in
Figure 5-1. The fact that a straight line log-log relation
is obtained for these data is not only an indication of
the invalidity of the theoretical semi-log penetration function,
it also indicat~s that the use of the inertial parameter may
satisfactorily account for the effect of the secondary vari-
ables.of particle and drop diameter and gas velocity. Thus,
there is reason to believe that a modified inertial impaction
theoretical treatment, coupled with a turbulent-mixing cap-
ture mechanism, may eventually supply the appropriate the-
oretical base for correct scrubber performance description.
However, from an engineering point of view, Equation (5-1)
is entirely adequate for most application or correlation pur-
poses for particulate removal by wet scrubbers, given a pre-
liminary evaluation of the' performance constant, K, with the
proper dust system.
F-93
-------
VI.
RESEARCH RECOMMENDATIONS
The correlations developed in this investigatio~;
as well as the successful differentiation of the primary
scrubber variables, provide an excellent base for t~esys-
tematic clarification of dispersed-drop wet scrubber cap-
abilities and limitations for particulate/hydrocarbon re-
moval from combustion process off-gases. It is recommended
that additional work be pursued on the following three sep-
arate levels, preferably concurrently, and within an in-
tegrated overall program:
1.
2.
Continuation of 'literature data correlation.
Development of a data-consistent theoretical;
mode 1 .
3.
Direct and relevant experimental testing.
A.
CONTINUATION OF LITERATURE DATA CORRELATION
This study made use largely of "primary" wet
scrubber literature data obtained by the direct experimental
testing of equipment in the laboratory or pilot plant. The
only industrial data utilized was that of Ingels (1960) on,
asphalt plants. Howev~r, there are many secondary reports
of scrubbing data and test results on full-scale plant equip-
ment in the trade journals, but because of time and budget
limitations, it was not possible to cover such a secondary
material. The first priority in any continuation of this
work should be the evaluation of these data in the context
of the correlating techniques already developed.
B.
DEVELOPMENT OF A SATISFACTORY THEORETICAL MODEL
The failure of the widely-accepted inertial im-
paction theory to give even a qualitative fit of the prime-
variable data for dispersed-liquid wet,scrubbers forces the
conclusion that this theory is an "emperor's clothes" sit-
uation. This theory does not cover the efficiency/scrubber-
variable relationships in its present form, and its apparent
successes in describing data do not bear close inspection.
While a.turbulent-agglomeration mechanism serves to provide
a rational basis for a gross qualitative interpretation of
F-94
-------
scrubber performance dependency on operating variables,
there are no quantitative predictive theoretical forms to
go ~long with this mechanism. Whether a suitable turbulent
agglomeration theory, modified impaction theory or a hybrid
model is to be emphasized is not clear at this point, but
any theory must (at minimum) explain the following points:
(a)
(b)
(c)
C.
The fundamental anomaly that, for a
given dust, penetration is inversely
proportional to the feed dust mass
concentration.
Inverse proportionality of penetration
to liquid-gas ratio and a power of pres-
sure drop.
Effect of dust and water particle size
and distribution on pe~etration.
EXPERIMENTAL PROGRAM
One of the major advantages of the correlation in
this study is that it describes the conditions necessary
for valid comparative rating of scrubber performance, and
allows a simple statement of the relative efficiency in the
form of the performance constant, K. While the influence
of the type of dust on performance constant remains to be
examined, the correlation allows testing under non-constant
conditions for a given dust without compromising the results.
A definitive laboratory experimental program is recommended,
utilizing two or three test scrubbers, and covering the fol-
lowing sub-programs:
(a)
(b)
(c)
Confirmation of the proposed per-
formance equations by full-range
testing of the prime variables of
dust loading, liquid-gas ratio,
and scrubber 9as velocity (or
pressure drop).
Direct verification of the indicated
influence of feed pre-agglomeration
on the feed dust loading/penetration
relationships.
Evaluation of the performance con-
stant, K, for the test scrubbers in
terms of nature and mean size of dust.
F-95
-------
( d )
(e)
(f)
Separation of the effects of liquid-dust
ag910meration from removal processes by
(1) independent control of the secondary
centrifugal removal force through spin-
out diameter variation, and (2) varying
inlet particle number and mass concen-
tration independently and measuring the
effect on penetration.
Separation of the liquid shearing phe-
nomenon by pre-atomization of the liquid
prior to injection into the scrubber.
Evaluation of the nature of condensation-
augmentation on scrubbing by pre-formed
fog injection in the absence qf condens-
ation.
F-96
-------
A, B
Cd
Cf
D
Dc
Dd
Dg
Do
Dp
Eo
E
Et
F
Fo
g
K
L/G
M
N
Nd
NOMENCLATURE
=
Cons tants
=
Dust concentration/unit gas volume, commonly
gr/CF
=
Drag coeffici~nt, dimensionless
Particle diameter
=
=
Count median diameter
=
Diameter of Water Droplet
Mass median diameter
=
=
Mean Sauter drop size, microns
=
Diameter of dust particle
Stage efficiency, constant
=
=
Overall scrubber efficiency, fractional
Target efficiency, dimensionless
=
=
Penetration, weight percent
Penetration of a single stage
=
=
Gravitational, constant
Proportionality constant
=
=
Ratio of volume flow of liquid to flow of gas at
vena contracta
=
Mass concentration
=
Water drop population concentration
Number concentrati'on: of drops, Cm-t~l:l
=
F-97
-------
Np = Number concentration of particles, cm~3
n = Number of stages
p = Press ure
Vg = Vel 0 city of gas
Vd = Vel oci ty of water drop
Vg = Relative air~water drop velocity, ft/sec (Vg . Vd)
WIG = Water/gas volumetric ratio, commonly gal/MCF
Xc = Agglomeration index, (Cd)(W/G)
a
=
Mass loss constant
B
=
Surface loss constant
~
=
Viscosity of gas
Liquid viscosity, poises
~L
PL
=
=
Liquid drop density
Density of dust particle
=
Pp
C1g
=
Geometric mean deviation
=
Liquid surface tension, dynes/em
C1L
'¥
=
Separation number, or inertial impaction parameter,
dimensionless
F~98
-------
(1)
( 2 )
( 3)
(4)
( 5)
( 6 )
(7)
(8 )
(9 )
( 10)
(11 )
( 12)
( 13)
BIBLIOGRAPHY
Accomazzo, M.A. and Nobe, K., Ind. Eng. Chem. (Prod.
Des. & Dev.) i, 425-430 (Oct. 1965).
American Industrial Hygiene Association, Air Pollution
Manual Part II, Control Equipment, Detroit, Michigan
~
Anon., "Catalytic Oxidation Controls Emissions", Env.
Sci. & Technology, 1, No. II, 1159 (Nov. 1969).
Anon., "Plant Licks Solvent Emission Problem", Env. Sci.
& Technology, i, No.2, 107 (Feb. 1970).
Barry, H.M., Chem. Eng. .li, 105-107, Feb. 8, 1960.
Boll ,. R.H., "Particle Collection and Pressure Drop in
Venturi Scrubbers", paper presented in the 69th Nat-
ion a 1 A I C h E Me e tin g, C i n c inn a t i, 0 hi 0, May 1 6, 1 971 .
Brewer, G.L., Chem. Eng., 75, No. 22, 160-165 (Oct. 14,
1968). . -
Brink, J.A., Jr. and Contant, C.E., Ind. Eng. Chem.,
~, 1157-60 (1958).
Byrd,"J.F. and Dewey, E.L., Chem. Eng. Prog., 53,
447-51 (1957). -
Calvert, S. & Legatski., L.K., "A Comprehensive State-
of~the-Art Evaluation for All Types of Dust Collection
Equipment that may be Applicable in Underground Coal
Mines", Contract No. S010023l, U.S. Bureau of Mines
Open File Report 4,..71, Dec. (1960).
Calvert, S. and Taheri, M., British Chem. Eng. .li, 254-
. 2 57 (1 9 6 6 ) .
Caretto, L.S. and Nobe, K., Ind. Eng. Chem., (Prod.
Des. & Dev.), i, 217-222 (July, 1966).
Cuffe, S.J. and Gerstle, R.W., "Emissions from Coal-
Fired Power Plants: A Comprehensive Summary", PB
174 70B,HEW, 1967.
F-99
-------
BIBLIOGRAPHY (continued)
( 14)
( 15)
( 16 )
( 17)
( 18)
( 19 )
(20)
(21 )
(22)
(23)
(24)
(25)
(26)
Dalla Valle, J.M., "Micromeritics", 2nd Ed., Putnam
Publishing Co~, (1948).
Danielson, J.A. (Ed.), "Air Pollution Engineering
Manual", HEW, PHS 999-AP 40, p. 206, (1967).
Davis, R.J. and Truitt, J., "The Function of Con-
densing Steam in Aerosol Scrubbers", Oak Ridge
National Laboratory Report No. 4654, March, 1971.
Drinker, P. and Hatch, T., "Industrial Dust", 2nd
Ed., McGraw-Hill Book Co., New York, New York (1954).
E 11 i 0 t t, J. , Kayne, N . an d LeDuc, M., "E x per 1 me n tal
Program for the Control of Organic Emissions from
Protective Coating Operations", Report No.8 Los
Angeles APCD, California, Jan. 1961.
Fahnoe, F., Lindroos, A.E. and Abelson, R.J., Ind.
Eng. Chern., 43, 1336-1346 (195l).
Fawcett, R.L., J. Air Poll. Control, 20, 461-465
(l970). -
First, M.W. et al, "Air Cleaning Studies" Progress
Report AEC Contract No. AT(30-1) 841, USAEC, NYO-
1586, Harvard Univ., (Feb., 1953).
Friedrich, H.E., J. Air Poll. Control Assoc., li,
924-928, (1969).
GCA Technology Division Preliminary Report, "Control
Techniques for Polycyclic Organic Matter Emission~",
HEW Contract No. CPA 70-43, August, 1970.
Golds(;hmid, Y. and Calvert, S., A.1oCh.E. Journal,
i, 352-358 (1963).
Hangebrauck, R.P., von Lehmeen, D.J. and Meeker,
J.R~, "Sources of Polynuclear Hydrocarbons in the
Atmosphere", PHS No. PB 174 706 (1967).
Hassler, J.W., "Actiyated Carbon", Chemical Publish-
ing Company, New York, New York, (1963).
F -100
-------
BIBLIOGRAPHY (continued)
(27)
(28)
(29)
( 30)
( 31)
(32)
(33)
(34)
(35)
(36)
(37)
(38)
. .
(39 )
(40)
HEW, PHS Publication No. 999-AP-40 (1967).
HEW, Pub. No. AP-51, "Control Techniques for Partic-
ulate Air Pollutants", NAPCA, January, 1969.
HEW, Pub. No. AP-68, "Control Techniques for Hydro-
carbon and Organic Solvent Emissions from Station-
ary Sources", NAPCA, March, 1970.
Ingels, R.M., Shaffer, N.R. & Danielson, J.A., J. Air
Po11 . Con t ro 1 As s 0 c., lQ, 29.. 33 (1960).
Johnson, G.A., Friedlander, S.K., Dennis. R., First,
M., and Silverman. L., Chem. Eng. Prog. li, 176-188
(1955).
Johnstone, H.F., Feild, R.B. and Tassler, M.C., Ind.
Eng. Chem., 46, 1601-1068 (1954).
Johnstone, H.F. and Robers, M.H.. Ind. Eng. Chem.
il, 2417-23, (1949).
Kerry, F.G. and Hu~i11, J.T., Chem. Eng. Progr..
il, (4). 37-41, (Apri 1, 1961).
Kleinschmidt, R.K., Chem. Met. Eng., .1.[, 487 (1939).
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Krista1, E., Dennis, R., and Silverman, L., J. Air
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Lancaster, B.W. and Strauss, W., IEC Fundamentals,
lQ, 362-368, August. 1971(a).
Lancaster, B.W. and Strauss, W., Chapter: Condensation
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Wiley-Interscience Div. of John Wiley & Sons, New York,
New Yo r k, 1971 (b ) .
Langmuir, I. and Blodgett, K.B., Report No. RL-225,
General Electric Research Lab., Schenectady, New
York, (1944).
F-101
-------
BIBLIOGRAPHY (continued)
( 41)
(42)
(43)
(44)
(45)
(46)
( 47)
(48)
(49)
(50)
( 51 )
(52)
(53)
(54)
Langstroth, G.O. and Gillespie, TOt Canadian Journal
of Re'search, fi, Sec. B, 45?-470, (1947).
Lapple, C.E. and Kamack, H.J., Chem~ Eng. Progress,
il, 110-121 (1955). ;
Levich, V., "Physiochemical Hydrodynamics", Prentice-
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Lunde, K.E. and Lapple, C.E., J. Air Poll~ Controll,
289-296, (1957).
Mantell, C.L., "Adsorption" 2nd Edition, McGraw-Hill,
New York, New York, 1961.
Matti a, M. M., Chern. Eng. Progr. &.£' No. 12, 24-79
(Dec. 1970).
Nukiyama, S. and Tanasawa, Y., Trans. SOC. Mech. Engr.
(Japan) i, No. 14, 86 (1938).
Ranz, W.E. and Wong, J.B., Ind. Eng. Chern., 44, 1371-81
(1952).
Sandomirsky, A.G. et al, "Fume Control in Rubber Pro-
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C.E. Ind. Eng. Chern. ~, 1615-1620 (1958).
Semrau, K.TOt J. Air Poll. Control AssocOt !Q., 200-207
(1960).
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(1963).
Silverman, L. and Davidson, R.A., J. Air Poll. Control,
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Stairmand, C.JOt J. Inst. Fuel (London), 30, 58-76"
February (1956).
F -1'02
-------
BIBLIOGRAPHY (continued)
(55)
(56)
(57)
(58)
Stern. A.C.. IIAir Pollution: Vol. III. Sources of Air
Pollution and Their Control II. Academic Press. New
York. New York 1968.
Walton. W.H. and Woolcock. A.. Intern. J. Air Pollution,
1, 26 (1960).
Werner, K.D.. Chern. Eng. 75, No. 24, 179-184 (Nov. 4.
1968). --
Wink1ep1eck, Water Sewage Works, Vol. 117:
1970.
R250-254,
F-l03
-------
APPENDIX G
QUESTIONNAIRE SURVEY ANALYSIS
-------
QUESTIONNAIRE SURVEY ANALYSIS
Discussion
. To complement the survey of the open literature and
related government contract reports, two questionnaire surveys
were conducted; one covering selected state and municipal or
regional control agencies and another covering selected trade
or industry associations.
In order to improve the chances of a reasonable re-
turn from these surveys, the questionnaires were kept to a
fairly simple format, primarily intended to ascertain if per-
tinent information had been developed in specific areas, such
as emission inventories, specific control procedures, legis-
lation in effect, cost of control and solid waste combustion
data. Copies of the two questionnaires and the distribution
lists for each are appended.
The control agency questionnaire was sent to 45
state and 33 municipal or regional agenci~s. The response
and type of data received as a result of the questionnaire
and follow-up contact are summarized in Table G-1.
Certain of the questions asked for specific data,
such as types or descriptions of problem pollutants, costs
of applied controls, and amounts of waste disposed of by
combustton. Responses to some of these questions are gtven
in Tables G-2 through G-4.
A number of control agencies indicated the avail-
ability of partial or incomplete unpublished-emission in-
ventory data if a personal visit were made to collect the
information. The questionable fruitfulness of such data
deterred us from pursuing personal visitation of this type.
This decision was based on:
(1) an assessment of published inventories
which indicated that most of the data
were based on estimates rather than de-
tailed surveys; and
G-1
-------
(2) phone and letter follow-up which gen-
erally confirmed the lack of specific
survey data.
The survey of trade and industry associations was
less fruitful than the control agency survey. Of the thirty-
eight associations contacted. only three provided any hard
data. Fifteen others responded. generally to indicate a
lack of information.
G-2
-------
G)
I
W
1.
2.
Information.
Type and Response
Questionnaire partially
comoleted
Emission Inventory Data
a. Partial estimates
provided
b. Detailed breakdown
~rovi ded
c. No data
3.
Legislative "Data (Controls !
or standards in effect) i
a. Some data provided i
b. No legislation in effect i
c. No data i
i
I
4.
Control Cost Data
a. Estimate orovided
b. No data
5.
Waste Combustion Data
a. Open burning
i. Total Estimate
ii. Detailed breakdown
1i1 Leaislative ban
iv. No data
b-. Inci nerati on
i. Total estimate
if. Detailed breakdown
111. No data
c. Agricultural and special!
waste burnings
i. Partial estimates
ii. No data
Table G-l - Control Agency Survey Summary
I State Agencies
I 45 surveyed. 33 responding
; No. of Agencies I % of Responding
! 30 i
! I
'3 .
i I
I 2: !
91
9
6
76
18
11
1
55
33
3
4
26
12
79
I
I
I
i~
, 3
I 61
I
!
. 1-
5
4
1
20
j
3
6
21
9
18
64
6
24
18
73
~fe~~:~~ndina ~
% of Responding Comments
73 ,i No ques t i onna i res were'
.. totally completed I
27 10nly 1 state and 3 :
I regions provided inven- L
I tories based on actual I
! survey data. Others range.
; from-guesstimates to
i limited source surveys.
14 regions and 11 states
: have control legislations.
; Others have_air quality i
!
i standards only.
Regional
33 surveyed.
No. of Agenci es
8
3
4
1
-
7
1
-
4
4
1
5
2
1
2
5
I
5
3-
36
9.
64
9
-
- j
I
36
36
;
9
45
18
9
18
45
9
I
1
I
I
45
27
-
I
_I 'Total estimates. rangi ng :
from wild guesses to ap-
Iparently valid approxima-
!tions. r
I !
lOne state and two regions'
: ban all open burning. ex- I
Icept by variance. Wide. I
:variations in quantities II
'burned. dependent on
estimating procedures. II
,Many states have 11ttle
__jor no municipal inciner- I
lation. Quantities estim-
!ated vary widely. I
iTwo regions and two
!states gave fairly de- j
itafled breakdown of '
lagricultural and forestry I
jbUrning.. . J
-------
Table G-2 - Specific Problem Pollutants Mentioned in Control
Agency Response to Question 16
Agency
California
Pollutants or Sources
Connecticut
Photochemically reactive hydrocarbon
Odorants - industrial
Idaho
Open burning of wastes, Tepee burners
automobile burning
Asphalt oil mists, Mercaptans, Phenols
Illinois
Kentucky
Mercaptans
M~rcaptans, Rendering plants
Maine
Maryland
Oregon
Drying ovens - can and bottle cap manu-
facture, Gasoline, Solvent recovery plant
o do rs
Organo-sulfur compounds from pulp mills,
Rendering plant odors
South Carolina
Ci ncf nnati
Mercaptans, Organo-phosphate odors
Solvents, Gasoline
Philadelphia
Amines, Mercaptans, Acrylics, Phthalic
anhydride
Mercaptans - pulp mills, Amines - render-
ing plants, Hydrocarbons - asphalt,
Terpenes - plywood driers, Resin manu-
facturing, Re-refining or destruction of
wastes
Puget Sound
San Francisco
Bay Area
Photochemically reactive organics,
Carbonyls from incineration
G-4
-------
Table G-3 - Control Costs Mentioned in Control Agency Response
to Questions 19 and 20
Agency
Mississippi
Question 19-
Do 11 a r Cos t of
Applied Controls
Question 20-
Dollar Cost of
Control for Total
Compliance
Oregon
South Carolina
3-1 /2~/person
$6,000,000
Cincinnati
$
$
400,000
500,000
$40,000,000
$22,000,000
$
500,000
Puget Sound
$8,000,000
$200,000,000
$25,000,000
$20,000,000
$400,000,000
Denver
San Francisco
Bay A re a
G-5
-------
Table
G-4
Haste
Combustion
Da ta
Provided
by
Control
Agencies
in
Response
to Question
28
G')
I
m
Control Open Burning Incineration
Agency Kun,- C omm- I ndul t- . House- Units or "unf - CODIID- Indust- House- ~'fi see 11 aneo us
cfpel erchl rhl hold Comments cfpel erchl rhl' hold Unfts
Arf zone 96.2 - -;.- ton/dey -' 2.9 .. - -> ton/dey 398 9.2 - ton/dey
slash ffres A9rfculture
Connectf cut none benned 4.057.000 - - none ton/yr 900 445 270 ton/yr
Agrf cuI tar. 0..011 tf on & 8ulty
I Constructfon Vastes
Oeh.... none none none .10.000 ton/yr none 20.000 50.000 none ton/yr none - ---. ....~
'eorgh - 163.000 18.200 444 .600 ton/yr 437,194 - - - ton/yr - - - -
Ideho 1.688 - - - tonl dey I - - - - - - - - -
New York (state) 210.000 32.000 21,000 789.000 ton/yr 1.671,000 119.000 119.000 477 ,000 ton/yr - - - -
Or.gon ..(..--..- -. 25,000. _.__....-...~_. ". .-....-.',. ton/yr - - 33.000 - ton/yr 29.000 17.000 - ton/yr
Vood lias tes Forest Shsh Fie 1 d 8ul'1lf n9
Pennsylvenh 3,000.000- ---...-> 1,700.000 Included ton/yr '1.000,000 , 300.000 Included ton/yr 500.000 - - ton/y r
wfth "un. wf tfl "un. Agrf culture
& Comm. & COIIII.
Rhode Ishnd ' 142.416 - - - ton/yr 135.180 233.611 ton/yr - - - -
'.
Vhconsln 75.000 - - 10.000 ton/yr 1,000,000 10,000 1.000 - ton/yr - - - -
Den v. r ..~- --.- .-- --- I-- none . ~...._. - -'-'-- - .__.~ benned none - .!- 4.378 - non. ~on/yr -- 1--- non. -- --- --- ---;..
-
Phfhdelphh .~----- I-- non. =l banned 604.000 270.000 - 330.000 ton/yr - - - -
.Puget Sound 28.200 155.000 78.000 ton/yr ' none 46.180 29. gOO 3.500 ton/yr 406.300 23.500 20.000 ton/yr
Sen Franchco 8a, non. non. none 2,200 ton/dey none. 400 100 none ton/dey 560 - - ton/dey
8anned after Agr1 cultUI'll
1969
-------
MSA RESEARCH CORPORATION
HYDROCARBON POLLUTANT SURVEY*
POtLUTION CONTROL AGENCY REPORT
OMB Clearance Number
85-F-70041
Expires 7-31-71
For the purposes of this ~tudYD the terms "hydrocarbon" and "organic" are
synonymous and include all organic compounds or classes of compounds from
stationary sources. If for som~ reason a partic~lar question cannot be
answered. write NOT AVAILABLE in the appropriate space. Additional entries.
information or comments may be ente~ed in Item No. 29 or o~ attached sheet.
If unpublished information is made available, results will be held
CONFIDENTIAL. .
Identification No.1
I For MSAR
Use Only
1 .
Agency
2.
Address
3.
Date
4.
Name of Person Completing Form
5.
Title
6.
Address & Phone No.
7. Pr1maryActivity in Pollution Control
o Ai r D Water D Soli d ~aste
8. Jurisdiction (State. County. etc. served by Agency)
Name of Jurisdictional Area ~oPulation Zoni ng or Use "as % of Total Area
Unit or Subdivision (so mi) thousands) Res.. Co mm . I n d. Aar. Other
9. If sources other than person designated in Item 4 are utilized in com-
pleting this form. indicate below the source and applicable item numbers.
Name, ..1 Title. I Organization I Item No.
10. .Has your agency conducted any surveys of hydrocarbon emissions. from
stationary sources? 0 Yes D No D Completed 0 In Prog.ress
11. Emission surveys completed .
Effective
Type Date Pub 11 shed Unpublished
Total organic emissions
Total organi c emissi ons by
area subdivision
°Spansored by Env1tonmental Protection Agency. Air Pollution Control Office.
Co~tract Number EHSO 11-12.
G-7
-------
Identification No.1
Emission surveys completed (continued)
Page 2 of 4
11.
Effective
Type Date Published Unpublished
Total organic emissions by
industry or operational
source
Specific compounds or classes
of emissions (describe)
If survey results are unpublished.
could be made available for study.
information is assured.)
E:nfS'Sion survey in progress
Type
check here 0 if resul ts
(Confidentiality of unpublished
; 2.
Estimated tompletion Date
13.
Check here [] if partially completed results could be made
available for study.
Has your agency compiled tables of hydrocarbon emissions for particular
industry or operational sources? [] Yes 0 No
Emission tables compiled
14.
Description
Published
Unpub li shed
If unpublished. check here
available for study.
o
if results could be made
15.
Are there specific organic air pollutants that
being particularly troublesome? 0 Yes
If yes, describe in Item 16.
Specific problem pollutants
you have identified as
o No
16.
Reason for Problem
Adverse Smog Lack of
T_vpe Health Effects Contribution Controls Other
"
..
- .. ~ --. -
, G-8
-------
17.
Identification No.1
Page 3 of 4
Has legislation been enacted to limit or control organic emissions
in your jurisdictional area? 0 Yes 0 No
If yes, describe briefly.
Legislative Time limit for
Leoislation Code No. ComDl1ance
18.
Estimate extent to which controls or limits have been applied
(Dercentaoe estimate for each cateaorv or leoislationL
l eo i s 1 at i ve 'C 0 n t r 01 Percent
19.
20.
Estimate dollar cost of currently applied controls.
$
Estimate dollar cost of control for total compliance.
$
21.
22.
23.
Does your agency have the responsibility for recommending control or
abatement procedures for specific emissions? DYes n No
If yes, have you developed a manual or handbook of contro~ngineering
procedures? 0 Yes 0 No
If possible, please forward copy of handbook.
Has your agency developed any economic evaluations (capital and
operating costs) of specific control procedures for organic emissions?
. 0 Yes n No
If yes, describe~riefly in Item 23.
24.
standards for any organic
Has your agency developed air quality
pollutants? DYes RNO
If yes, describe briefly 1n tern 25.
Ai l't t d d .
. r Qua 1 ;v s an ar s "
Comoound or class Concentration 11 mi ts
25
G-9
-------
Identification No.1
Page 40f 4
26.
Has your agency developed emission standards (or limits) for specific
sources of organi c emi ssi ons1 n Yes 0 No
If .yes. describe briefly in Item ~.
27.
Type of emission
Emission limits
28.
Estimate tonnage of solid waste disposal by combustion for each of the
following categories.
Category Tonnage
Open Burning
MuniciDal dumD or landfill .
Commercial and Institutional
Industrial
Household
Incineration
MuniciDal
Commercial and TnsutTtutfona I
Industnal
tfousehold
Other (Agricultural. etc.)
29. . Continuation items
Item no. Additional information
.
..
..
G-10
-------
HYDROCARBON POLLUTANT SURVEY*
TRADE ASSOCIATION REPORT
OMB Clearance Number
. 85-S-70043
"Expires 5-31-71
MSA RESEARCH CORPORATION
For the purpose of this study, the terms "hydrocarbon" and "organic"
are synonymous and include all organic compounds or classes of com-
pounds from stationary sources. .
Identification NoJ
I~~~ ~~~;
Date
2.
Address
3.
o.
7.
If sources other than person designated in Item 4 are utilized in completing
this form, indicate below the source and applicable item numbers.
Name
Title
Oraanization
I tern No.
8. Pri nci Da 1 industrvlies) represented b' association
No. Companies % of Total
Des.c ri Dt ion SIC No. ReDresented Industrv
9.
Has your association conducted any industry surveys of organic emissions?
DYes 0 No 0 Completed 0 In Progress
10.
Emission surveys completed
Effective
Tvpe Date Pub li shed Unpublished
Total Industry Emissions
Emissions by Specific
Operational Source
~Sponsored by Environmental Protection "Agency, Air Pollution Control Office,
Contract No. EHSD 71-12.
G-ll
-------
Identification No.1
Page 2 of 3
10. Emission surveys comDleted (continued)
Effective
Tvoe Date Publ1shed Unoubl1shed
Particular Compounds
or Classes (Describe)
Other (Descr1be)
If survey results are unpublished, check here [] if results could be
made available for confidential study.
11.
Emission surveys in progress
Type
Estimated Completion Date
Check here D if partially completed survey results could be" made
available for confidential study.
12.
Are there specific organic emissions that you have identified as being
known or potential pollutant problems? DYes D No
If yes, describe briefly
13.
Have you developed any manuals or handbooks of air pollution control
procedures? 0 Yes 0 No
If
d
"b
yes. es cn e
DescriDtion Published Unoub1ished
If unpublished, check here
confidential study.
[] if data could be made available for
G-12
-------
Identification No.1
Page 3 of 3
14.
Have you developed any economic ev~luations (capital
costs) of specific control procedures? . [J Yes
If es describe
Descr1 tion
and opera t 1 ng .
D No
If unpublished, check here
for confidential study.
[] if results could be made available
15.
Are there particular areas of research and development relating to
control of hydrocarbon emissions that you feel should be pursued?
DYes [] No
If yes, describe briefly
16.
Continuation Items, Additional Information or Comments
G-13
-------
DISTRIBUTION LIST FOR CONTROL AGENCY SURVEY
Arthur N. Beck
Director, Bureau of Environmental
Department of Public' Health
State Office Building
Montgomery, Alabama 36104
James E. Fibbe.
Deputy Director of Bureau
Mobile County Board of Health
248 Cox Street
Mobile, Alabama 36604
Health
James A. Anderegg
Chief, Branch of Environmental Health
Alaska Department of Health and Welfare
Po uch H
Juneau, Alaska 99801
Norman E. Schell
Director, Air Pollution Control Division
Arizona State Department of Health
4019 North 33rd Avenue
Phoenix, Arizona 85017
. Robert C. Taylor
Director, Bureau of Air Sanitation
Maricopa County Health Department
1825 East Roosevelt
Phoenix, Arizona 85006
S. Ladd Davies, Director
Arkansas Pollution Control
1100 Harrington Avenue
Little Rock, Arkansas 72202
Commission
John A. Maga '
Executive Officer
California Air Resources Board
1400 10th Street
Sacramento, California 95814
G-14
-------
D.J. Callaghan
Chief Administrative Officer
Bay Area Air Pollution Control
939 Ellis Street
San Francisco, California 94109
Louis J. Fuller
Air Pollution Control Officer
Los Angeles County Air Pollution
434 South San Pedro Street
L~s Angeles, California 90013
District
Control District
Joseph Palomba, Jr.
Chief, Air Pollution Control
Division of Air, Occupational
Colorado Department of Public
4210 East 11th Avenue
Denver, Colorado
and Radiation Hygiene
Health
Fred H. Longenberger
Chief Air Pollution Control Engineer
Denver Air Pollution Control Agency
1445 Cleveland Pl.
Denver, Colorado 80202
Louis Proulx, Jr.
Section Chief
Air Pollution Control Section
Environmental Health Services Division
Connecticut State Department of Health
79 Elm Street
Hartford, Connecticut 06115
James 1. Wilburn
Director, Air Pollution Control Division
Delaware Water and Air Resources Commission
Post Office Box #916
Dover, Delaware 19901
Charles Couchman
Chief, Division of Air Pollution Control
District of Columbia Department of Public
1875 Connecticut Avenue
Washington, D.C. 20009
He a 1 th
G-15
-------
Sidney A. Berkowitz
Director, Bureau of Sanitary Engineering
State Board of Health
Florida Air and Water Pollution Control Commission
Suite 400
315 South Calhoun Street
Tallahassee, Florida 32301
Paul William Leach, Director
Metropolitan Dade County Pollution
864 N.W. 23rd Street
Miami, Florida 33127
William A. Hansell
Director, Air Quality Control Branch
Georgia Department of Public Health
47 Trinity Avenue, S.W.
Atlanta, Georgia 30334
Cont~ol Department
B.I. Garland
Chief, Industrial Hygiene and Air
Fulton County Health Department
99 Butler Street, S.E.
Atlanta, Georgia 30303
Vaughn Anderson, Director
Air Pollution Control Section
Idaho Department of Health
512 West State
Boise, Idaho 83707
Pollution Control
Robert R. French
Chief, Bureau of Air Pollution
Illinois Air Pollution Control
616 State Office Building
Springfield, Illinois 62706
Control
Board
William J. Stanley, Director
City of Chicago
Department of Air Pollution
320 North Clark Street
Chicago, Illinois 60610
Harry D. Williams
Director, Division of Air Pollution
Indiana Air Pollution Control Board
1330 West Michigan Stre~t
Indianapolis, Indiana 46206
Control
Control
G-16
-------
Lewis F. Scott, Director
Bureau of Air Pollution Control
Room 1642
City-County Building
Indianapolis, Indiana 46204
Paul H. Houser
Chief, Environmental Engineering
Iowa State Department of Health
Lucas State Office Building
Des Moines, Iowa 50319
Service
James Clark
Kansas Air Quality Conservation
State Office Building
Topeka, Kansas 66612
John A. Noon, Jr.
Air Pollution Control
Air Pollution Control
Kansas City-Wyandotte
1014 Armstrong Street
Kansas City, Kansas 66102
Commission
Engineer
Division
County Health Department
Ralph C. Pickard
Executive Secretary
Kentucky Air Pollution Control Commission
275 East Main Street
Frankfort, Kentucky 40601
Ralph Bourne
Chief Engineer
Louisville and Jefferson County
Control District
621 West Jefferson Avenue
Louisville, Kentucky 40202
Air Pollution
Vernon C. Parker
Chief, Air Control Section
Louisiana Air Control Commission
Post Office Box #60630
New Orleans, Louisiana 70160
Raeburn Macdonald
Water and Air Environmental
State of Maine
Au~usta, Maine 04330
Improvement Commission
6-17
-------
Jean J. Schueneman .
Chief, Division of Air Quality Control
Maryland State Department of Health
2305 North Charles
Baltimore, Maryland 21218
Elkins W. Dahle, Jr.
Director, Bureau of Industrial Hygiene
Division of Air Pollution Control
Bureau of Industrial Hygiene
Baltimore City Health Department
602 American Building
Baltimore, Maryland 21202
John C. Collins
Chief, Bureau of Environmental Sanitation
Division of Air Pollution and Radiological
Department of Public Health
600 Washington Street
Boston, Massachusetts 02111
Frank J. Reinhardt
District Director
Metropolitan Air Pollution
600 Washington Street
Boston, Massachusetts 02111
Control District
B.D. Bloomfield
Chief, Air Pollution Control Section
Division of Occupational Health
Michigan Department of Public Health
3500 North Logan
Lansing, Michigan 48914
Morton Sterling, Director
Air Pollution Control Division
Wayne County Department of Health
414 City-County Building
Two,Woodward Avenue
Detroit, Michigan 48226
Edward M. Wiik
Director, Division of Air Quality
Minnesota Pollution Control Agency
Health Building
University of Minnesota Campus
Minneapolis, Minnesota 55440
G-18
Health
-------
APPENDIX H
BREAKDOWN OF FUEL CONSUMPTION
AND ESTIMATED EMISSIONS
AND
REVIEW OF ELEMENTARY COMBUSTION STUDIES
-------
1.
Breakdown of Fuel Consumption and Estimated Emissions
In Tables H-l through H-10 are presented state
and regional breakdowns of fuel consumption and estimated
hydrocarbon emissions for each of the major fos~il fuels~
These data amplify the discussion and summary presented in
Volume I, Section III.
Data sources for the fuel consumption by area
and consuming sector were the Federal Power Commission (1968)
and the u.S. Bureau of Mines (1968, 1970).
2.
Review of Elementary Combustion Studies
In the course of the literature review on hydro-
carbon emissions from combustion sources, it was found
that very few studies have been made of the trace organic
constituents from combustion under field or full-scale con-
ditions. Most of the reported studies have treate~ only
"total gaseous hydrocarbon" measured by flame ionization or,
in a few cases, reported determination of chemical groups,
such as aldehydes (or total oxygenates), organic acids,
unsaturates, and the like. As an attempt to gain some in-
sight into the types and amounts of particular chemical
species emitted from combustion sutides reported in the'
literature were briefly reviewed.
Combustion processes are basically chemical re-
action between a fuel and an oxidant at more or less ele-
vated temperatures. A generalization about such reactions
is that the reaction complexity increases as the chemical
complexity of the reactants increases. Thus, our ~nder-
standing of the reactions which take place is based pri-
marily on analogies drawn from studies of the simpler
systems. No attempt was made to review all of the liter-
ature on combustion. For good general 'reviews, reference
may be made to such texts as Lewis and von Elbe (1961) and
Lewis, Pease and Taylor (1956).
The oxtdation of simple hydrocarbons is generally
viewed as a free radical initiated, chain branching mech-
anism, with the critical intermediates being partially
oxygenated species containing peroxy, hydroperoxy, and
aldehyde linkages. Newman and Gal (1968) have summarized
the sequence of elementary steps in the oxidation of
methane by the following scheme:
H-l
-------
CH4
t
(CH300)
t
CHfoH
HCHO )ro CO
t
(HCO)
k
(HCOO) . C02
Although the concentrations and total number of
stable intermediates and final combustion products vary with
the conditions of the particular studys the consensus of
investigators of elementary combustion studies i~ that the
most common species formed are formaldehydes acetaldehyde,
formic acid, acrolein and acetone. Reference may be made to
Bonner and Tipper (1965) and Pease (1934s 1935) as typical
studies leading to this conclusion.
Actual combustion processes are complicated by
the fact that most fuels (and their combustion intermediates)
can undergo thermal degradations cracking or condensation
reactions which do not evolve direct oxication. Thuss
studies of thermal decomposition in the absence or insuffi-
ciency of oxygen must be included in our consideration of
potential contaminant emissions. Examples of such s~udies
are the work of Holmes and Shaw (1961), Roberts and Clough
(1963), and Martin (1965)s on the thermal decomposition of
cellulosic materials. Again, the stable intermediates iden-
tified were the lower a1dehydess ketoness and acids. In
addition to the oxygenated species, the low molecular weight
hydrocarbons, principally methanes ethane and ethy1enes
have been identified as stable products.
H-2
-------
Most of the fundamental studies of combustion have,
in order to allow menaingful experimental design and inter-
pretation, been limited to simple hydrocarbon fuels. When
one considers the potential variety of combustible material
that is involved in the combustion of fuels and wastes, the
extrapolation of the results of fundamental studies becomes
quite tenuous. The breakdown of types and concentrations of
emission products from combustion thus depends on specific
studies of the particular fuel or waste composition and the
combustion process used. Such specific studies have not been
made, except for limited special cases, such as current
studies of combustion of high plastic content wastes.
H-3
-------
Table H-l - State Breakdown of Coal Consumption by Electric
Utilities (1968)
Emissions
Coal Consumed Hydrocarbon (tons)
Rank State (103 tons)- (as CH4)- Aldehydes
,
1 Ohio 29,569 4,435 74
2 Illinois 27,123 3,918 68
3 Pennsylvania 25,160 3,774 63
4 Indiana 21,745 3,263 54.
5 Michigan 18,983 2,847 47
6 North Carolina 15,056 2,259 38
7 Alabama 14,368 2,156 36
8 Kentucky 13,255 1 ,989 33
9 New York 13,152 1,973 3~
10 West Virginia 13,021 1,953 33
11 Tennessee 12,835 1,926 32
12 Virginia 7,869 1, 181 20
13 Wisconsin 7,626 1 , 145 19
14 Maryl and 7,447 1,118 19
15 Missouri 7,421 1,113 19
16 Georgia 7,118 1 ,068 18
17 New Jersey 5,025 755 13
18 Florida 4,481 672 11
19 r~i nnesota 4,205 632 11
20 Iowa 2,803 420 7
21 Colorado 2,742 411 7
22 North Dakota 2,732 410 7
23 Massachusetts 2,730 410 7
24 Arizona 2,722 410 7
25 Connecticut 2,585 389 6
26 Wyoming 2,085 314 5
27 South Carolina 1 , 822 273 5
28 Delaware 1 ,368 206 3
29 New Hampshire 765 116 2
30 District of Columbia 725 110 2
31 Nevada 536 81 1
32 Kansas 461 69 1
33 Nebraska 419 63 1
34 Utah 396 60 1
35 Mississippi 241 36 1
36 Rhode Island 110 17
37 Vermont 32 5 -
-
Totals 280,733 41,977 704
H-4
-------
T~ble H-2 - Regional And State Breakdown of Non-Utility Consumption
of Coal, 1968(a)
Region & State
New Eng land
Massachusetts
Connecticut
Me.. N. H.. V t.. & R. I .
Middle Atlantic
New York
New Jersey
Pennsylvania
E. North Central
Ohio
Indiana
l111no1s
Michigan
Wisconsin
W. North Central
Minnesota
Iowa
M1ssouri
North & South Dakota
Nebraska & Kansas
South At lant i c
Delaware & Maryland
District of Columbia
V i rg I n Ia
West Virginia
North Carolina
South Carol1na
Georgia & Florida
East S. Central
Kentucky
Tennessee
Alabama & Mississippi
West S. Central
Ark., la., Okla., & Texas
Mountain
Colorado
Utah
Montana & Idaho
Wyoming
New Mexico
Arizona & Nevada
Pacific
Washington & Oregon
California
Alaska
Totals
Coal Consumed
103 tons
Industrial (b)
Emissions {ton/yrJ
Hydrocarbons Aldehydes
337
346
141
6,7211
878
9,247
14,949
6,203
8,863
9,567
4,418
1,123
1,788
1,932
341
214
1.046
131
4,740
5,505
2,335
1,159
573
2,326
2,852
1,896
115
501
338
216
234
16
16
306
14
~
92,029
(a) Bureau of Mines data
(b) 2xcludes coke and gas plants
169
173
71
3,364
439
4,624
7,475
3,102
4,432
4,784
2,209
562
894
966
171
107
523
66
2,370'
2,753
1,168
580
287
1,163
1,426
948
58
251
169
108
117
8
8
153
7
--.ill
46,023
H-5
0.8
0.9
0.4
16.8
2.2
23.1
37.4
15.5
22.2
23.9
11.0
2.8
4.5
4.8
0.9
0.5
2.6
0.3
11.9
13.8
5.8
2.9
1.4
5.8
7.1
4.7
0.3
1.3
0.8
0.5
0.6
0.8
..J.:!
229.9
Domestic & Commercial
Coal Consumed Emissions {ton/yrJ
103 tons Hydrocarbons Aldehydes
114
1
22
153
7
699
171
33
0.3
0.4
1.8
230
11
1,049
3,050
1,497
4,968
2,211
2,993
2,033
998
3,312
1,474
1,995
795
263
149
312
51
5.1
2.5
8.3
3.7
5.0
1.3
0.7
0.4
0.8
0.1
1,193
395
224
468
77
135
125
1,080
638
929
413
308
0.2
0.2
1.8
1.1
1.6
0.7
0.5
1.4
1.5
0.3
90
83
720
425
619
275
205
660
584
105
840
876
158
20
13
316
134
351
30
4
18
474
201
527
45
6
27
215
17
-M
0.8
0.3
0.9
0.1
143
11
~
17,101
0.4
....Q.J.
25,670
42.3
-------
Table H-3 - State Breakdown of Natural Gas Consumption by
, Electric Utilities (1968)
Gas E m is's ion s, (tons)
Consumed Total
Rank S ta te (106 cu ft) Hydrocarbons Aldehydes
1 Texas 731,359 1 ,462 1 ,098
2 California 674,713 1 ,349 1,011
3 Louisiana 290,920 582 435
4 Oklahoma 147,642 295 222
5 Florida 127,881 256 192
6 Kansas 1 20 , 805 242 180
7 New York 97,849 196 147
8 Mississippi 76,496 153 114
9 Arkansas 72, 142 144 108
10 Illinois 6 L 751 124 93
11 Iowa 54,810 110 81
1 2 Minnesota 44,599 89 66
1 3 Arizona 41,171 82 63
14 Nebraska 29,343 59 45
15 Colorado 27,466 55 42
16 South Carolina 27, 1 42 54 42
1 7 New Mexico 23,965 48 36
18 Missouri 23,330 47 36
19 Tennessee 22,034 44 ~3
20 Ne\'l J e rs ey 18,263 37 27
21 Wisconsin 1 7,883 36 27
22 Georgia, 14,944 30 21
23 Nevada 14,740 30 21
24 Alabama 11 ,467 23 18
25 Ohio 6,629 1 3 9
26 Indiana 5,635 11 9
27 Utah 4,224 8 6
28 Massachusetts 3,477 7 6
29 Pennsylvania 3,048 6 6
30 South Dakota 1 ,670 3 3
Totals 2,797,398 5,595 4,197
H-6
-------
TABLE H-A - REGIONAL AND STATE BREAKDOUN OF NATURAL GAS CONSU~'PTION IN 1968(a)
INDUSTRIAL USAGE(b)
Reqion & State
New England
Connecticut
Me, N.H., Vt.
Massachusetts
Rhode Island
Middle Atlantic
New Jersey
New York
Pennsylvania
East North Central
Illinois
Indiana
Michigan
Ohio
\~isconsin
West North Central
Iowa
Kansas
Minnesota
Missouri
Nebraska
North Dakota
South Dakota
South Atlantic
De 1 a\'/a re
Florida
Georgia
Maryland and D. C.
North Carolina
South Carolina
Virginia
West Virginia
East South Central
Alabama
Kentucky
Mississippi
Tennessee
West South Central
Arkansas
Louisiana
Oklahoma
Texas
Mountain
Arizona
Colorado
Idaho
Montana
Nevada
New Mexico
Utah
Wyoming
Pacific
Alaska
California
Ore go n
~Jashi ngton
Gas Consumption
MMcf
Emiisions (tons/yr)
Hydrocarbons Aldehydes
15,045
1,387
22,379
5 , 188
53
5
78
18
248
321
1 ,102
23
2
34
8
106
138
472
70,850
91,679
314,913
336,393
231,137
250,906
350,281
117,315
1.177
809
879
1,226
411
289
515
310
330
169
9
14
505
346
376
525
176
124
221
133
141
72
4
6
82,646
147,183
88,518
94,287
48,166
2,612
4,101
11 ,377
88,299
140,115
37,514
60,538
68,641
42,493
88,069
40
309
490
131
212
240
149
308
601
182
419
376
17
132
210
56
91
103
64
132
258
78
179
161
171,755
51,969
119,590
107.476
138,997
864,848
118,948
1,696,231
487
3,027
416
5,937
209
1 ,297
178
2,544
54,613
83.683
24,637
23,155
8,693
76,864
53.639
34,794
2,677
544,292
51,755
93,370
191
293
86
81
30
269
188
122
9
1 ,905
181
327
82
126
37
35
13
115
80
52
4
816
78
140
Totals
7,134,018
24,969
10,699
(a) Bureau of Mines data
(b) excludes utilities, but includes refiflery fuel use
H-7
-------
Table H-5 - Regional ~nd State Breakdown of Natural Gas Consumption
in 1968 a , Residential and Commercial arid Institutional
Usage
Resfdentlal Commercfal and Instftutfonal
Gas Consumed Emfssfons (tons/yr) Gas Consumed Emissions (tons/yr)
Regfon & State MMcf Hydrocarbons Aldehydes MMcf Hydrocarbons Aldehydes
New England
Connectfcut 26,437 13 133 11,224 6 56
Me., N.H., Vt. 3,571 2 18 1,672 1 9
Massachusetts 74,919 38 375 25,396 13 128
Rhode Is land 10,605 5 53 3,411 2 18
Mfddle Atlantfc
New Jersey 137,116 69 685 32,546 16 163
New York 319,282 160 1,596 122,885 62 615
Pennsylvania 285,978 143 1,430 87,620 44 438
E. North Central
l11fno15 392,325 196 1,961 174,565 87 873
Indiana 145,955 73 730 60,661 30 304
Mfchfgan 315,694 158 1,579 117,124 59 585
Ohfo 444,964 223 2,225 165,414 83 828
W15cons f n 93,425 47 466 36,067 18 180
W. North Central
Iowa 84,936 43 425 48,034 24 240
Kansas 89,372 45 446 46,232 23 231
Mfnnesota 90,410 45 453 65,536 33 328
M15sourl 138,764 69 694 79,821 40 399
Nebraska 53,376 27 268 41,765 21 209
North Oakota 7,169 4 36 7,072 4 35
South Oakota 10,302 5 51 10,723 5 54
South Atlantfc
Oeleware 7,068 4 35 2,084 1 10
Florida 11,318 6 56 21 ,890 11 110
Georgia 84,072 42 420 36,034 18 180
Maryland & D.C. 79,015 40 395 30,419 15 153
North Carolfna 24,646 12 124 20,624 10 103
South Carolfne 16,756 8 84 10,544 5 53
Vf rgf nfa 43,582 22 218 24,594 12 123
West Vfrgfnfa 54,665 27 274 20,402 10 103
E. South Central
Alabama 51 ,708 26 259 34,749 17 174
Kentucky 75,824 38 379 36,089 18 180
M15s15sfppf 29,526 15 148 18,297 9 91
Tennessee 43,784 22 219 38,325 19 191
W. South Central
Arkansas 56,346 28 283 37 ,886 19 190
Louisfana 77,762 39 389 56,937 29 285
Oklahoma 74,782 37 374 42,751 21 214
Texas 211,763 106 1,058 139,442 70 698
Mountain
Arizona 26,681 13 134 23,389 12 118
Colorado 78,371 39 391 47,287 24 236
Idaho 6,545 3 33 6,374 3 31
Montana 19,711 10 98 13,651 7 69
Nevada 5,493 3 28 6,997 4 35
New Mexfco 31 ,568 16 158 30,713 15 154
Utah 40,779 20 204 8,114 4 40
Wyoming 12,592 6 63 11,637 6 59
Paciffc
Alaska 2,293 1 11 4,713 2 24
Ca If forn fa 517,636 259 2,588 189,903 95 950
Oregon 15,126 8 76 7,874 4 39
Washfngton 26,342 --..ll 131 16.244 8 --'!!.
Tota 1 s 4,450,354 2,229 22,254 2,075,736 1,035 10,387
H-8
-------
Table H-6 - Consumption and Emissionf From Fuel Oil Usage by Electric
Utilities
Electric utility company use
(Thousand barrels) (1) 1969 Emissions
P. A. D. District Distillate-type oils Residual-type oils (Tons/year)
and S ta te 1969 l.2.M. 2) 1969 ill.!!. Total lIydrocarbon Aldehyde
District 1:
Connecticut 331 28 18,215 12,772 1,947 389
Delaware 85 56 1,086 454 123 25
District of Columbia 3,236 1,432 340 68
Florida 290 314 35,686 32 .261 3.759 752
Georgia 180 176 493 406 71 14
Maine 96 115 4,727 4,986 506 101
Maryland 408 238 4,143 1.440 478 96
r~assachusetts 370 217 31 ,490 23,450 3.345 669
New Hampshire 3 3 2.092 1.305 220 44
New Jersey 565 430 34.028 26.643 3,632 726
New York 561 475 43,565 34,971 4.633 927
North Carolina 828 447 693 174 160 32
Pennsylvania 2,073 1,399 20.557 10.527 2.376 475
Rhode Island 17 36 3.152 2,426 333 67
South Carolina 490 205 1.582 680 218 44
Vermont 108 19 130 169 25 5
Virginia 486 99 9,350 362 1.033 207
West Virginia ~ --.ll 196 131 ~ ---1
Total 6,920 4,279 214,421 155.089 23,223 4,646
District 2:
Illinois 494 183 1,171 515 175 35
Indiana 195 165 271 223 49 10
Iowa 258 229 60 69 33 7
Kansas 171 168 188 264 38 8
Kentucky 135 119 35 12 18 4
Michigan 320 269 1,147 920 154 31
Minnesota 416 262 752 568 123 25
Missouri 123 122 76 57 21 4
Nebraska 107 86 78 46 19 4
North Dakota 80 66 11 18 10 2
Ohio 467 450 397 328 91 18
Ok !ahoma 42 38 6 24 5 1
South Dakota 64 54 78 25 15 3
Tennessee 3 3
Wiscons i n -lil. -1i 175 187 ---11 -.!
Total 3,112 2.248 4,445 3,256 794 161
District 3:
Alabama 142 138 5 15 3
Arkansas 12 9 318 149 35 7
Lou is 1a na 79 44 13 21 10 2
Mississippi 1 1 289 100 30 6
New Mexico 9 9 33 26 4 1
Texas 1,487 L.ili --1! ---ll -1!l. -1!
Total 1.730 1.646 704 330 255 51
District 4:
Colorado 19 21 233 103 26 5
Idaho 15 7 2
Montana 14 7 105 23 12 2
Utah 25 21 1,619 1,432 173 35
Wyoming -1.!!. --1!. --i. -----1l ~
Total 91 72 1,966 1,579 216 42
District 5:
Alaska 207 177 195 162 42 8
Arizona 6 7 40 43 5 1
California 29 28 19,761 18,978 2,078 416
Hawa if 52 38 6.029 5~439 639 128
Nevada 11 14 51 58 7 1
Oregon 8 10 1
Washington - --1! --.11. --1
Total 305 264 26,098 24.702 2.773 554
United States. tots 1 12,158 8,509 247,634 184,956 27,261 5,454
!1! Source: Federal Power Commission
2 Revised to include data for gas-turbine plants
H-9
-------
Tab.le H-7 - Consumption and Emissions from Industrial (Excluding Oil
Companies) Usage of Fuel Oil
Industrial Use
(Thousand barrels) 1969 Emissions
P.A.D. District Distillate fuel 0115 Residual fuel 0115 (Tons/year)
and State 1969 1968 1969 .ill..!!. Total Hydrocarbon Aldehyde
District 1:
Connecticut 513 905 9,411 10,941 625 219
Delaware 146 153 3,087 2,602 204 71
District of Columbia 4 5 104 63 7 2
Florida 1,041 1,057 5,424 4,905 407 158
Georgia 929 1,005 5,743 6,184 420 160
Maine 371 373 1,999 1,261 149 58
Maryland 743 833 2,225 2,323 187 78
Massachusetts 833 957 10,316 10,764 703 252
New Hampshire 63 115 1,203 594 80 28
New Jersey 1,935 1,904 13,304 14,009 960 361
New York 3,125 4,086 16,055 18,496' 1,208 468
North Carolina 818 1,059 3,563 2,910 276 109
Penn sylvania 2,489 3,226 9,981 12,730 786 314
Rhode Is land 245 238 1,953 1,722 138 51
South Carolina 405 398 1,104 1,477 95 40
Vermont 116 197 395 235 32 13
VI rgl nia 939 997 1,178 1,590 133 64
West Virginia -1l.Q. --ili ~ ..LM!. --1.l2. ..-1L
Total 15,090 17,941 88,561 94,369 6,529 2,489
District 2:
Illinois 1,838 2,11 5 5,139 6,179 440 185
I nd lana 1,169 1,206 2,503 4,620 231 102
Iowa 245 257 42 51 18 11
Kansas 110 120 419 354 33 13
Kentucky 660 605 707 599 86 43
Mlchl9an 1,706 2,044 2,804 2,307 284 130
Minnesota 910 1,015 2,428 2,144 210 89
MI ssourl 532 559 1,223 518 111 48
Nebraska 47 67 249 130 19 7
North Dakota 145 100 57 19 13 7
Ohio 2,937 2,936 3,434 3,339 401 195
Oklahoma 143 207 349 334 31 13
South Dakota 89 72 33 26 8 4
Tennessee 584 676 758 276 85 40
Wisconsin ~ -.-.1J.i -lli. ...!....W. ---2.! --!!
Total 11 ,800 12,693 20,952 21,901 2,064 933
District 3:
Alabama 826 790 1,483 817 145 66
Arkansas 212 194 136 455 22 12
louis iana 797 773 994 433 113 54
Mississippi 528 466 62 106 37 23
New Mexico 249 210 4 7 16 11
Texas ~ ...hill ---1R ~ --..ill. -Ii
Total 4,155 3,681 3,131 2,578 459 240
District 4:
Colorado 226 763 1,075 748 82 32
Idaho 518 621 166 114 43 25
Montana 148 114 232 102 24 11
Utah 436 493 3,127 3,081 224 84
Wyoming --ill. --1Ql ---1.l!! -ill -il ---1Q
Total 1,645 2,192 4,938 4,280 414 172
District 5:
Alaska 384 375 681 571 67 30
Arizona 1,476 1,345 70 16 97 63
California 5,116 4,766 7,710 7.,690 808 377
Hawaii 228 162 524 515 47 21
Nevada 324 305 2 20 14
Oregon 1,122 1 .047 1.778 1,436 183 84
Washington ~ --L.ll!!. ..L.lli ~ .J.li --.n
Tota 1 9.766 9,288 13.072 12,536 1,438 684
United States, total 42,456 45,795 130,654 135.664 10.904 4,518
H-10
-------
Table H-8 - Consumption and Emissions from Oil Company Usage of
Fuel on
Oil Company Use 1969 Emissions
(Thousand barrels) type oils (1)
P.A.O. Oistrict Distillate type oil s Residual (Tons/year)
and State 1969 ill.!!. ill! ill.!!. Total Hydrocarbon Aldehydes
District 1:
Connecticut 269 48 77 80 22 13
Delaware 57 6 122 125 11 5
District of Columbia 10 1 5 4 1
Florida 349 152 247 205 38 20
Georgia 66 17 3 16 4 3
Maine 49 7 14 17 4 2
Maryland 281 45 392 462 42 20
Massachusetts 104 42 79 203 12 6
New Hampshire 1 2 3 3
New Jersey 221 64 6,344 6,681 414 143
New York 120 73 67 68 12 6
North Carolina 86 82 46 43 8 5
Pennsy1v6n!a 1,761 1,408 4,743 5,524 410 174
Rhode Island 62 13 11 13 5 3
South Carolina 72 34 102 100 12 5
Vermont 11 2 4 1 1
Virginia 106 67 185 147 18 8
West Vlrghia -1.Q.l. ~ --12. --1. ---2
Tota 1 3,726 2,079 12,460 13,692 l,009 418
Oistrict 2: 823(2)
111 Inois 1,818 3,268 2,448 320 145
Indiana 1,877 1,200 3,831 4,101 360 159
Iowa 30 16 2 1
Kansas 29 16 797 794 52 18
Kentucky 22 33 50 50 4 2
Michigan 274 159 851 818 71 29
Minnesota 29 71 870 887 57 19
Missouri 24 51 469 493 31 11
Nebraska 9 4
North Dakota 10 9 372 472 24 8
Ohio 302 95 929 1 ,161 78 32
Oklahoma 16 15 197 406 13 5
South Oakota 27 18 2 1
Tennessee 15 34 2 10 1
Wiscons I n -1J.i .--2! --M ----ill. ---1! ---1.Q.
Total 4,697 2,638(2) 11 ,700 11 ,756 1,033 402
Oistrlct 3:
Alabama 11 18 41 42 3 1
Arkansas 9 15 27 28 2 1
Lou is !ana 940 864 21 19 61 40
Miss iss I ppl 7 15 131 135 9 3
New Mexico 11 13 3 7 1
Texas 2,164 1,898 ~ ...lJill. -1.2l --1.l!
Total 3,142 2,823 1,243 1,248 277 157
District 4:
Colorado 128 10 1 4 8 5
Idaho 5 6 1
Montana 60 10 683 681 47 17
Utah 115 83 400 420 32 13
Wyoming -ill -lli -----ill. --ill. --M --1l
Total 859 631 1,567 1,592 152 68
Oistrict 5:
Alas ka 163 150 16 16 11 7
Arizona 17 16 1
California 514 1,004 6,550 8,098 445 129
Hawaii 13 12 666 679 43 15
Nevada 111 108 7 5
Oregon 214 92 324 305 34 16
Washington -ill ....Jl1 2,033 1,943 --ill. -!..Q.
Total 1,443 1,804 9,589 11 ,041 696 232
United States, total 13,867 9,975(2) 36,559 39,329 3,167 1,277
!1~ Includes a small amount of crude oil
2 Revised
H-ll
-------
Table H-9 - Consumption and Emissions from Commercial and Institutional
Usage of Fuel 0i1*
. Heating OIls 1969 Emhslons
(Thousand Barrels) (T onsl yea r)
P.A.O. Dhtrlct 110. 5 No.6 Total Total
1ITId 'S1a te U69 1968' 1969 1968 1969 1968 Hydrocarbon Aldehydes
Dhtrlct 1:
ConnectIcut 34 227 4.706 3,630 4,740 3.857 299 100
Oehwa re 195 135 228 181 423 316 27 9
Dfstrlct of Columbfa 505 762 7.161 6.841 7.666 7.603 483 161
florida 100 51 400 745 500 796 32 11
Georgia 1.542 597 806 529 2.348 1.126 148 49
Ma Ine 385 141 1.582 1.177 1.967 1.318 124 41
~rylend 1.301 1.383 3.817 3.134 5,118 4.517 322 107
Massachusetts 7.544 7.498 22.129 24,277 29,673 31.775 1.870 617
New HampshIre 382 120 635 322 1.017 442 64 21
New Jersey 1.170 1.102 11.659 11.912 12.829 13.014 808 267
New York 1.994 2.718 57.533 57,524 59".527 60.242 3.750 1.238
North Caro 11 na 579 552 503 432 1.082 984 68 22
Pennsylvania 4.941 5.635 6.447 6.294 11 .388 11.929 717 239
Rhode I s lend 420 325 1.367 1.204 1.787 1.529 113 38
South CarolIna 247 220 58 147 305 367 19 6
¥~'f"tMtft t 18 39 123 81 141 120 9 3
Vlr91nfa 430 653 1.304 1.893 1.734 2.546 109 36
West Ylrglnfa JlQ --1.l1 --1! --li -12.!. ---.ill. ----11 -..!
Total 21.957 22,271 120.479 120,370 142.436 142.641 8.974 2.969
District 2:
illinois 1.~26 6.896 7.138 6.910 14,664 13.806 924 308
'1 ndlana 524 329 1.786 1.355 2.310 1,684 145 48
Iowa 142 154 61 66 203 220 13 4
Kansas 10 33 16 46 26 79 2 1
Kentucky 6 4 97 17 103 21 6 2
MIchIgan 422 365 833 480 1,255 845 79 26
MInnesota 593 119 421 78 1,014 197 64 31
Hfssourl 1.004 1,367 640 670 1.644 2.037 104 35
Nebraska 62 84 73 124 135 208 9 3
North Dakota 10 4 10 2 20 6 1
Ohio 101 124 328 249 429 373 27 9
Ok lahoma 2 5 10 7 10
South Dakota 33 3 26 15 59 18 4
Tennessee 10 25 10 25 1
Wfscons I n -ill --.!1!1. --lli --ill ~ ~ -.!l ---1.!.
Total 11,124 9.937 11 .758 10.372 22.882 20.309 1.442 489
District 3:
Alabama 74 74 74 74 5 2
Arkansas
Louisiana 6 6
MississIppI 50 50 100 6 2
New MexIco 2 4 2 4
Texas --..ll --U -2.Q1 ---1.!i -ill. --.111 -1f ----11
Total 199 143 559 114 758 257 47 16
Dfstrlct 4:
Colorado 223 215 223 112 446 327 28 9
Idaho 62 31 103 47 165 78 10 3
Montana 200 284 26 135 226 419 14 5
Utah 389 277 373 288 762 565 48 16
Wyoming --1li --1.Q.Q. -ill -ill -2.!Q --1Q1 -11. -LL
Total 1.089 907 1.020 785 2.109 1.692 132 44
District 5:
Alaska 3 19 13 22 13 1
ArIzona 19 19 16 16 35 35 2 1
California 942 938 1.495 1.490 2.437 2.428 154 51
Hawaii 110 101 159 148 269 249 17 6
Nevada 27 30 24 29 51 59 3 1
Oregon 2.060 1.868 1.340 1.135 3.400 3.003 214 71
Washington 2.3U 2.292 1.369 1.348 3.696 3.640 -ill ---1!
Tota I 5.488 5.248 4.422 4.179 9.910 9..427 624 208
United .S~~t~~. ~ota' 39.95' 38.506 138.238 135.820 178.095 174.326 11.219 3.726
*May Include some usage as IndustrIal heatIng
H-12
-------
Table H-10 -
Consumption and Emissions
from Domestic Usage of Fuel
0i1*
rlo. 1 Heating Oi15
(Thousand Barrels) 1 969 Emiss ions
r----. 1969 (Tons/year)
?A.O. District Automatic Other Total
and State Burners Heating ~ No.4 Tota 1 Hydrocarbons Aldehydes
District 1:
Connecticut 219 141 16.896 1.525 18.781 1.183 789
De lawa re 245 128 2.891 10 3.274 206 137
District of Columbia 74 400 2.090 165 2.729 172 115
Florida 101 560 3.178 78 3.917 247 165
Georgia . 89 111 1.919 2.119 133 89
Maine 262 636 8.009 208 9.115 574 383
Maryland 23B 57 11.775 182 12.252 772 515
Massachusetts 961 337 49.461 2.183 52.942 3.335 2.223
New Hampshire 170 183 6.637 45 7.035 443 295
New Jersey 384 150 40.361 7.828 48.723 3.070 2.047
New York 2.045 684 79.449 10.360 92.538 5.830 3.887
North Carol1na 723 568 10.614 51 11.956 753 502
Pennsylvania 800 252 37.932 1.725 40.709 2.565 1.710
Rhode Island 37 105 7.068 100 7.310 461 307
South Carolina 266 266 3.761 100 4.393 277 185
Vermont 154 67 4.644 83 4.948 312 208
V~rgfnia 527 255 10.261 254 11 .297 712 475
West Virginia --.! -1. ----ill ~ --!.Qi ~ --ll
Total 1969 7.299 4.902 297.539 24.903 334.643 21.083 14.057
Tota 1 1968 8.557 5.873 291.759 23.456
District 2:
Illinois 2.316 695 19.429 952 23.392 1.474 983
I ndhna 1.727 1.022 14.068 250 17 .067 1.075 717
Iowa 1.453 487 5.204 16 7.160 451 301
Kansas 106 18 609 11 744 47 31
Kentucky. 251 67 1.333 32 1.683 106 71
Michigan 4.943 1.869 19.347 301 26.460 1.667 1.111
Minnesota 2.380 983 10.914 154 14.431 909 606
Missouri 537 438 4.918 222 6.115 385 257
Nebraska 285 107 1.578 1.970 124 83
North Dakota 730 133 2.663 14 3.540 223 149
Ohio 1.427 185 13.398 239 15.249 961 641
Oklahoma 99 91 529 719 45 30
South Dakota 632 124 1.477 2,233 141 94
Tennessee 49 65 1 ,801 1.915 121 81
Wisconsin 2.457 -ill 1 5 .883 ---iiL 19.840 1.250 --!.li
Tota 1 1969 19.392 7.242 113.151 2.733 142,518 8,979 5,988
Total 1968 21.143 8.083 115.617 2.367
District 3:
Alabama 18 8 301 327 21 14
Arkansas 28 14 346 388 24 16
Louisiana 32 19 606 657 41 27
Miss iss i ppi 26 30 671 727 46 31
New Mexico 8 30 187 225 14 9
Texas -lli -..ill ...L.Q.1Q ----:..L ~ -1.!!1. -L!.
Total 1969 525 341 3.151 3 4.020 253 168
Total 1968 513 312 1.650 6
District 4:
Colorado 355 257 617 1,229 77 51
Idaho 876 689 1.708 3 3.276 206 137
Montano 89 564 544 1,197 75 . 50
Utah 92 364 667 32 1.155 73 49
Wyoming ~ -1L ----EQ --.l --111. --1i -1.Q.
Total 1969 1.426 1.911 4.206 36 7.579 476 317
Total 1968 1.797 1 .615 5.646 50
District 5:
Alaska 848 1.660 2.508 158 105
Arhona 62 115 177 11 7
Cal1fornfa 293 947 1.240 78 52
Hawaii 19 109 128 8 5
Nevada 95 376 471 30 20
Oregon 1.958 4.827 6.785 427 285
Washfngton 3.387 --!hlli 11 .699 -1.ll .-m.
Total 1969 6.662 16.346 23.008 1.449 965
Total 1968 6.302 15.936
United States total. 1969 35.304 14.396 434.393 27.675 511.768 32.240 21.495
Unfted States total. 1968 38.312 15.883 430.608 25.879
*May include some oomaercfal usage
H-13
-------
REFERENCES
Bone» W.A. and Hill, G.Sq "S10w Combustion of Ethane"
Proc. Roy. Soc. (London) A129, 434-57 (1930).
Bone, W.A. and Gardner, J.B., "Comparative Studies in the
Slow Combustion of Methane, Methyl Alcohol, Formaldehyde
and Formic Acid", Proc. Roy. Soc. (London) A154, 297-328
Bonner, B.H. and Tipper, C.F.Hq "Coo1-F1ame Combustion of
Hydrocarbons", Tenth S m osium International on Com-
bustion", pp. 145- 50, Combustlon Instltute, 9 5
Federal Power Commission, "Steam-E1ectric Plant Construction
Cost and Annual Production Expenses", Twenty-first Annual
Supplement, 1968
Holmes, F.H. and Shaw, C.J.G., liThe Pyrolysis of Cellulose
and the Action of Flame Retardant, I. Significance and
Analysis of the Tar", J. App1. Chem. li, 210-16 (1961).
Lewis, B., Pease, R.N. and Taylor H.S., CombustiDn Processes,
Princeton University Press, 1956
Lewis, B. and von E1be, G., Combus~ion, Flames and Explosions
of Gases, Second Edition, Academic Press, New York, 1961
Martin, S.,
eria1s b,¥
national)
1965
"Diffusion-Controlled Ignition of Cellulosic Mat-
Intense Radiant Energy", Tenth Symposium (Inter-
on Combustion, pp 877-896, CombustiDn Institute,
Newman, r~.Bo and Get1, D., liOn the Sequence of Elementary Steps
in Gas Phase Hydrocarbon Oxidation", Comb. Flame, 12, 371-9
(1968). -
Pease, R.N., liThe Mechanism of the Slow Oxidation of Propane",
J. Am. Chem. Soc. iI, 2296-9 (1935)
Pea s e, R . N. . and M u n r 0, W. P .» II The S low 0 x i d a t ion 0 f Pro pan e",
J. Am. Chem. Soc. 56, 2035-8 (1934)
Roberts, A.F. and Clough, G., "Therma1 Decomposition of Wood
in a n I n e r tAt m 0 s p her e", N i nth S y m p 0 s i u m (I n t ern a t ion a 1) 0 n
Combustion, pp 158-66, Combustion Institute, 1963
U.S. Bureau of Mines, Mineral Industry Surveys, "Shipments of
Fuel Oi 1 and Kerosene", 1970
U.S. Bureau of Mines, Minerals Yearbook, 1968
H-14
-------
APPENDIX I
WASTE COMBUSTION EMISSION FACTORS
AND
MUNICIPAL WASTE BREAKDOWN
-------
1.
Waste Combustion Emission Factors
Introduction
The estimation of hydrocarbon emissions from solid
waste combustion detailed in Volume I, Section III, required
th~ application of selected emission factors for the various
waste combustion modes. This appendix summarizes the pub-
lished emission factors utilized in the development of t~e
waste combustion emission estimates.
Discussion
A review of the literature on solid waste combustion
emission disclosed a variety of emission factors, ranging from
those derived from analytical measurem~nt of the emissions
from full scale operations to estimates and extrapolations
from limited small scale test data. Many of the reported values
are estimates developed by the author of the particular pub-
lication from his evaluation of prior literature and, partic-
ularly in the case of those reported by Duprey (1968) and
McGraw and Duprey (1971), include unpublished data and private
communications.
In summarizing
Table 1-1, the source or
ing the reported factors
"Comments".
the reported values, as shown in
type of information used in develop-
are noted under the column headed
The selection of average or "typical" emission
factors from the wide range of reported values was highly
arbitrary. Not only are the analytical measurements sup-
porting the reported values poorly documented, but, in ad-
dition, little attempt was made to classify the type or com-
position of the waste being burned. The wide variety of
wastes handled and the extreme ranges of actual combustion
conditions makes assignment of 'typical' factors quite spec-
ulative.
In selecting the values shown in Table 1-2, used
for our emission estimates, an attempt was made to give greater
weight to those data based on full-scale or reasonably simu-
lated test conditions. Also weighted heavily were those
studies which attempted to obtain detailed component break-
downs. Much of the reported emission data is based on measure-
ment of total gaseous hydrocarbons by flame iQnization an-
alysis and thus does not generally include the contribution
of such species as aldehydes and organic acid~. Those few
1-1
-------
Table
1-1
Suml11ary
of Emission
Factors
for Was te
Combustion
-
I
N
m 55 on ac ors. n S
Combustion Mode b t.ota 1 Total Saturated Unsatu ra ted ! Or9an1 c Aldehydes PAH Co-ents. ll8fe rences
roanl cs Hydrocarbons Hydroca rbons Hydrocarbons i Acids
I
l'Iunlc1pal Incineration 1.5 ! Consensus of 11 te ra tu re IIcGraw & Duprey (1971)
(multiple chamber) I
i data
0.3 Consensus of 11 terature Duprey (1968)
data
0.8 0.15 0.1 0.05 0.3 Review of 8ay Area Feldstein, et al (1963)
studies
1.58 (ave) ; Simulated full-scale Rose, et al (1959)
test measurement<
0.3 - 6.5 Simulated full-scale Stenbur9, et al (1960, 1961'
test measurements
0.9 - 6.3 0.06 - 0.16 0.001-0.84 Full scale test data A. D. II ttl e (1970)
:
2.7 0.005 Consensus of literature A. D. little (1970)
and test data
Domestic Inctneratton 2 With primary air -- McGraw & Duprey (1971)
(sln91e chamber) .- ltterature consensus
100 Without primary air -- McGraw & Duprey (1971)
literature conSAnCUC
252 Emission factor used. for Say Area APCD (1969)
. 'restdenthl burnlnQ"
1 - 2 2'- 7, 2.5-5.5 Consensus of literature Duprey (1968)
data
250 30 36 24 35 Rework of YOCOII (1956). Feldstein, et al (1963)
.' I "Backyard & S1n91e Chamber"
20 Estimated from Iltera- . Ho vey, eta I (1966)
ture data
Industrhl and Commercial 3 Multiple Chamber -- McGraw & Duprey (1971)
r' Inctneratlon'.. literature consensus
...
15 S1n91e chamber -- McGraw & Duprey (1971)
Itterature consensus
20 Conical burners -- McGraw & Duprey (1971)
lIunlclpal refuse
11 Conical burners u McGraw & Duprey (1971)
wood wutes
-
0.005 Full scale test data Hangebrauck, et al (1967)
-0.029
90 'Stn9le chillber" Bay Area APCD (1969)
em15slon factor
0.8 , "Double chamber" and Bay Area APCD (1969)
; "Silo" wood burner
I I
0.5-0.8, 3 0.2-1 literature Consensus -- Duprey (19U)
lIultlp1e chlmber
2 6-25 2-3 l fterature consensus u Duprey (1968)
f1 ue fed
.. '.
- -.. -
E t
F
ib Iton of wlste burned
-------
Table
1-1
(continued)
.......
I
c..,J
COl:1bustlon Mode Total Total Saturated Unsaturated Organic Aldehydes PAH
Organics Hydrocarbons Hydrocarbons Hydrocarbons Acids Comments Reference
Open Burning 30 I!unlclpal Refuse -- McGraw a Duprey (1971)
literature consensus
4-20 Ag rl cu ltural Refuse -- RcGraw a Duprey (1971)
literature consensus
30 Auto components -- McGraw a Ouprey (1971)
literature consensus
.-
30 Municipal Refuse -- Duprey (1968)
literature consensus
12 13 0.01 Landscape and agricultural Duprey (1968)
literature consensus
30 30-40S of 14-16 0.01 0.007 Munl cl pal Refuse -- Gerstle and Kemnitz (1967)
total hvdrocarbon simulated tes t data
-
30 30-40S of 8-18 0.005 0.OD7 !..andscape refuse -- Gerstle and Kemnitz (1967)
total hydrocarbon simulated test date
250 Agricultural burning Bay Area APCA (1969)
emission factor
252 30 36 24 35 Land clearing -- rework Feldstein, et al (1963)
of Yocom (1956)
Simulated test data -
2-36 0.5-2.1 0.3-8.7 Darley (1966)
agrl cultural refuse
.. -
4-19 0.4-3.3 1.0-5.6 Simulated test data -- Boubel (1969)
112.3 ave.) grass and crop stubble
9-15 Field test data -- Boubel (1969)
(10.6 ave.) grass stubble
80 40 1-3 Domestic refuse -- estlm- Hovey, et al (1966)
ated from literature
14-15 Municipal refuse -- .Hangebrauck et al (1967)
, test data
-------
detailed studies reported indicate that the contribution of
partially oxygenated species to the total organic emissions
can be quite significant, exceeding in some cases the 'total
hydrocarbon' emission.
Much study remains to be done in order to arrive
at more meaningful emission estimates from solid waste com-
bustion. However, it is quite clear that alternatives to
open burning and indiscriminate use of poorly designed arid
uncontrolled incinerators must be implemented.
2.
Municipal Waste Breakdown
Introduction
In the development of estimates of emissions from
solid waste combustion, detailed in Chapter III, Part C, a
state-by-state breakdown of urban waste generation and dis-
posal was made. This appendix presents the results and dis-
cussion of this breakdown.
Discussion
In the Systems Study of Air Pollution from Municipal
Incinerators (PB 192378, A.D. Little, Inc., 1970r-urban waste
generation factors were presented for each state, based on an
average generation rate of 5 1b/person day for household
wastes plus a seasonally adjusted factor for yard wastes.
Using the A.D. Little generation factors and population data
from the U.S~ Bureau of the Census, state estimates of total
urban waste generation were compiled, as shown in Table I-3.
By this procedure, the total urban waste generation for the
United States is approximately 1.73 x 108 tons/year.
In order to estimate the amount of generated urban
waste that is burned, some additional assumptions were made.
That portion of the generated waste handled by municipal in-
cineration was estimated from the assumption that municipal
incinerators operate at fifty percent of listed capacity.
State data on incinerator capacity were obtained from the
A.D. Little report and from the National Emission Standards
Study (Report to Congress of the Secretary of Health, Ed~
ucation and Welfare, March, 1970). Open burning was assumed
to be the disposal mode for 25 percent of the total generated
waste.
I-4
-------
TAB L E .. I - 2
SELECTED EMISSION FACTORS
FOR SOLID WASTE COMBUSTION SOURCES
Combustion Mode
Total Organic Emissions
(lb/ton burned)-
Incineration
Municipal
Industrial and Commercial
Conical Burners (wood)
Domestic
Open Burning
Municipal
Industrial and Commercial
Lumbering
Field and Crop
Land Clearance
Demolition and Construction
Coal Refuse
Forest Wildfires
1-5
4
14
20
40
45
30
30
20
40
30
20
40
-------
Exceptions to the above assumptions were made where
the questionnaire survey returns from the pollution control
agencies provided data on the amounts of solid waste disposal
by combustion, or where other data sources reported signifi-
cant population areas having prohibitions on open burning.
The estimates summarized in Table I-3 thus present
a crude approximation of the magnitude of the urban waste
generation and disposal by combustion on a state-by-state
basis.
I-6
-------
TAB LEI - 3
STATE BREAKDOWN OF ESTIMATES
OF WASTE GENERATION AND COMBUSTION
Waste Urban Has te
Generation State Urban Wastes Combustion
(lb/person Population Fraction (l0 5 ton/ (105 ton/
S ta te day) (1,000) (%) yr) y r)
Alabama 6.4 3,373 60 23.638 A
B 5.91
Arizona 6~5 1 ,752 91 18.911 A
B 0.4*
Arkansas 6.2 1 ,886 49 10.451 A
B 2.61
California 6.8 19,715 94 229.984 A
B 57.50
Colorado 5.8 2,178 85 19.597 A
B 4.90
Connecticut 6.0 2,988 87 28.,465 A 9.00
B 0.01*
Delaware 6.0 543 75 4.460 A
B 1. 12
Florida 6.8 6,671 86 71.197 A 8.05
B 2.44*
Georgia 6,2 4,492 65 33.036 A 4.37*
B 4.45*
Idaho 5.9 698 55 4.135 A
B 6.16*
Illinois 6.0 10,978 86 103,379 A 6.02
B 25.85
Indiana 6.0 5,143 65 36.606 A 1. 33
B 9.15
Iowa 5.9 2,790 55 16.524 A
B 4.13
Kansas 6.0 2,222 65 15.815 A 0.06
B 3.95
1..7
-------
TABLE 1-3 (continued)
Waste Urban Was te
Generation State Urban Was tes Comb~stion
(lb/person Population Fraction (105 ton/ (105 ton/
State day) (1,000) (%) yr) y r)
Kentucky 6. 1 3,161 48 16.892 A 2. 10
B 4.22
Louisiana 6.4 3,564 69 30.547 A 2.92
B 7.63
Maine 5.8 977 60 6.205 A
B 1. 55
Maryland 6.2 3,875 83 36.391 A 3.93
B 1 .6*
Massachusetts 5.9 5,630 87 52.739 A 7. 15
B 13.18
Michigan 5.9 8,778 79 74.668 A 5. 15
B 18.67
Minnesota 5.8 3,768 67 26.722 A 0.42
B 6.68
Mississippi 6.2 2,159 43 10.505 A 0.25
B 2.63
Missouri 6. 1 4,636 71 36.642 A 1. 52
B 9. 16
Montana 5.7 682 55 3.902 A
B 0.98
Nebraska 5.9 1,468 60 7.658 A
B 1. 91
Nevada 5.7 482 91 4.563 A 0.25
B 1. 14
New Hampshire 5.7 732 67 5.041 A 0.22
B 1. 26
New Jersey 6.0 7,085 97 75.070 A 2.25
B 18.77
New Mexico 6.0 998 73 7.979 A
B 2.00
1 - 8
-------
TABLE 1-3 (continued)
Haste Urban ~1 a s t e
Generation S tate Urban Wastes Combustion
(lb/person Population Fraction {10s ton/ (1()5 ton/
State day) (ltOOO) (%) yr) yr)
New York 6.2 1 7 t 9 80 91 185.135 A 31.65
B 9.99*
North Carolina 6.2 4t962 45 25.265 A
B 6.32
North Dakota 5.7 611 37 2.351 A
B 0.59
Ohio 6.0 10t542 78 90.038 A 7.35
B 22.51
Oklahoma 6. 1 2t498 69 19.188 A
B 4.80
Oregon 6.3 2t056 69 16.312 A O. 10
B 0.25*
Pennsylvania 6.2 11 t 670 74 97.714 A 10.0
B 30.0*
Rhode Island 6.0 922 89 8.986 A 1. 35
B 1.42*
South Carolina 6.2 2t523 45 12.848 A
B 3.21
South Dakota 5.7 661 44 3.026 A
B 0.76
Tennessee 6.2 3t839 57 24.758 A
B 6. 19
Texas 6.3 10t989 86 108.657 A 1. 42
B 27.17
Utah 6.0 lt061 86 9.994 A 0.52
B 2.50
Vermont 5.7 438 40 1 .825. A
B 0.46
Virginia 6.3 4t686 69 37.175 A 2.64
B 9.30
1-9
-------
TABLE 1..3 (continued)
Waste Urban ~Jaste
Generation State Urban Was tes Combustion
(1 b/pers on Population Fraction (105 ton! (105 ton/
S tate day) (1,000) (%) yr) yr)
Washington 6.4 3,353 72 28.204 A 0.52
B 1. 77
Wisconsin 5.9 4,367 69 32.445 A 13 *
13 8. 11
Wyoming 5.6 329 57 1.916 A
B 0.48
Dist. of Columbia 6.2 746 100 8.442 A 2.8
B
Totals 200,251 1,733,082
A = Incineration, B = Open Burning, * = Data from questionnaire
1-10
-------
~
REFERENCES FOR TABLE 1-1
A. ~~c~~~~~~i o~~~~e s~~1~m~: ,A~H~J 1 ~~~~n 7~:o~0 M~
Bay Area Air Pollution Control District, Source Inventory
of Air Pollutant Emissions (1969).
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